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WELDING ENGINEERING

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Welding and joining processes

Process terminology

The European standard, BS EN ISO 4063:2010 Welding and allied processes - Nomenclature of processes and reference numbers, assigns a unique number to the main welding processes. These are grouped as follows:

Arc welding Resistance welding Gas welding Forge welding Other welding processes Brazing, soldering and braze weldingEach process is identified within the group by a numerical index or reference number. For example, the MIG welding process has a reference number of 131 which is derived as follows:

1 - Arc welding 3 - Gas-shielded metal arc welding 1 - Metal arc inert gas weldingThe main arc welding process reference numbers are:

111 manual metal arc welding; 114 self-shielded tubular-cored arc welding; 121 submerged arc welding with one wire electrode; 125 submerged arc welding with tubular cored electrode; 131 metal inert gas welding (MIG welding); 135 metal active gas welding (MAG welding); 136 tubular cored metal arc welding with active gas shield; 141 tungsten inert gas arc welding (TIG welding); 15 plasma arc welding;The reference numbers are used as a convenient way of identifying the welding process in documentation such as welding procedures (BS EN ISO 15614 series) and welder qualification (BS EN 287 and BS EN 9606 series) records.

Process options

Factors which must be taken into account when choosing a suitable welding or joining process are:

material type product form: plate or tubular quality and strength requirements degree of mechanisation capital costAlthough consideration of these factors will identify the most suitable welding process, the choice within a company may be restricted by the cost of implementing a new process, availability of plant or current workforce skill. Welding and joining processes available to the welding engineer can be separated into the following generic types:

Fusiono arco gaso power beam

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o resistance Thermomechanicalo frictiono flasho explosive

Mechanicalo fasteners

Solid stateo adhesiveo solderingo brazing

The suitability of the processes for welding and joining materials, joint types and components are shown in Table 1.

Table 1. Suitability of the processes for welding and joining materials, joint types and components

Process Index no. Steel Stainless Al Butt

jointLap joint Plate Tube Portability Manual Mechanised

Automated Site

Arc 1 Yes Yes Yes Yes Yes Yes Yes Yes Yes Yes YesGas 3 Yes Possible Possible Yes Yes Yes Yes Yes Yes No YesLaser 52 Yes Yes Possible Yes Yes Yes Yes No No Yes NoResistance 2 Yes Yes Yes Possible Yes Yes Possible Possible Yes Yes NoFriction 42 Yes Yes Yes Yes No Yes No No No Yes NoBrazing 9 Yes Yes Yes No Yes Yes Possible Yes Yes Possible YesFasteners none Yes Yes Yes No Yes Yes No Possible Yes Yes YesAdhesives none Yes Yes Yes No Yes Yes Yes Yes Yes Possible YesIn selecting a suitable process, consideration must also be given to the type of application, for example, the portability of equipment, whether it can be used on site, whether it is manual or mechanised, and the overall cost of the welding plant.

Fusion welding processes

When welding using a fusion process, the edges of a component are melted together to form weld metal.

Table 2. Heat source, mode of shielding, thickness range and metal deposition rates for various fusion processes

Process Heat source ShieldParent metal 

thickness mm

Deposition rate kg/hr

ArcMMA Arc Gas/flux 1-100 1-4MIG Arc Gas 0.5-100 1-8TIG Arc Gas 0.1-100 1-4SAW Arc Flux 5-250 5-20ES/EG Resistance/arc Flux/gas 5-250 5-20Stud Arc - 4-20 -

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GasOxyfuel Flame Gas 0.6-10 1-2

Power beamLaser Radiation Gas 0.2-25 -EB Electrons Vacuum 0.2-250 -

ResistanceSpot/Seam Resistance - 0.2-10 -

ThermitThermit Chemical Gas 10-100 -Table 2 shows heat source, mode of shielding, thickness range and metal deposition rates for a range of fusion processes. Although fusion welding is one of the simplest joining techniques, problems likely to occur include porosity in the weld metal, and cracking in either the weld or heat affected zone (HAZ). Porosity is avoided by ensuring adequate shielding of the weld pool and, for materials such as aluminium, the addition of filler wire.

Consideration of the joint design and the chemistry of the weld metal will prevent weld metal cracking. HAZ cracking which might be caused by hydrogen, is avoided by using low hydrogen consumables and controlling the heat input and the rate of cooling of the parent metal.

The Manual Metal Arc process (MMA Welding)

Manual metal arc welding was first invented in Russia in 1888. It involved a bare metal rod with no flux coating to give a protective gas shield. The development of coated electrodes did not occur until the early 1900s when the Kjellberg process was invented in Sweden and the Quasi-arc method was introduced in the UK. It is worth noting that coated electrodes were slow to be adopted because of their high cost. However, it was inevitable that as the demand for sound welds grew, manual metal arc became synonymous with coated electrodes. When an arc is struck between the metal rod (electrode) and the workpiece, both the rod and workpiece surface melt to form a weld pool. Simultaneous melting of the flux coating on the rod will form gas and slag which protects the weld pool from the surrounding atmosphere. The slag will solidify and cool and must be chipped off the weld bead once the weld run is complete (or before the next weld pass is deposited).

The process allows only short lengths of weld to be produced before a new electrode needs to be inserted in the holder. Weld penetration is low and the quality of the weld deposit is highly dependent on the skill of the welder.

Types of flux/electrodes

Arc stability, depth of penetration, metal deposition rate and positional capability are greatly influenced by the chemical composition of the flux coating on the electrode. Electrodes can be divided into three main groups:

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Cellulosic Rutile BasicCellulosic electrodes contain a high proportion of cellulose in the coating and are characterised by a deeply penetrating arc and a rapid burn-off rate giving high welding speeds. Weld deposit can be coarse and with fluid slag, deslagging can be difficult. These electrodes are easy to use in any position and are noted for their use in the 'stovepipe' welding technique.

Features:

deep penetration in all positions suitability for vertical down welding reasonably good mechanical properties high level of hydrogen generated - risk of cracking in the heat affected zone (HAZ)Rutile electrodes contain a high proportion of titanium oxide (rutile) in the coating. Titanium oxide promotes easy arc ignition, smooth arc operation and low spatter. These electrodes are general purpose electrodes with good welding properties. They can be used with AC and DC power sources and in all positions. The electrodes are especially suitable for welding fillet joints in the horizontal/vertical (H/V) position.

Features:

moderate weld metal mechanical properties good bead profile produced through the viscous slag positional welding possible with a fluid slag (containing fluoride) easily removable slagBasic electrodes contain a high proportion of calcium carbonate (limestone) and calcium fluoride (fluorspar) in the coating. This makes their slag coating more fluid than rutile coatings - this is also fast-freezing which assists welding in the vertical and overhead position. These electrodes are used for welding medium and heavy section fabrications where higher weld quality, good mechanical properties and resistance to cracking (due to high restraint) are required.

Features:

low hydrogen weld metal requires high welding currents/speeds poor bead profile (convex and coarse surface profile) slag removal difficultMetal powder electrodes contain an addition of metal powder to the flux coating to increase the maximum permissible welding current level. Thus, for a given electrode size, the metal deposition rate and efficiency (percentage of the metal deposited) are increased compared with an electrode containing no iron powder in the coating. The slag is normally easily removed. Iron powder electrodes are mainly used in the flat and H/V positions to take advantage of the higher deposition rates. Efficiencies as high as 130 to 140% can be achieved for rutile and basic electrodes without marked deterioration of the arcing characteristics but the arc tends to be less forceful which reduces bead penetration.

Power source

Electrodes can be operated with AC and DC power supplies. Not all DC electrodes can be operated on AC power sources, however AC electrodes may be used on either AC or DC.

Welding current

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Welding current level is determined by the size of electrode - the normal operating range and current are recommended by manufacturers. Typical operating ranges for a selection of electrode sizes are illustrated in the table. As a rule of thumb when selecting a suitable current level, an electrode will require about 40A per millimetre (diameter). Therefore, the preferred current level for a 4mm diameter electrode would be 160A, but the acceptable operating range is 140 to 180A.

What's new

Transistor (inverter) technology is now enabling very small and comparatively low weight power sources to be produced. These power sources are finding increasing use for site welding where they can be readily transported from job to job. As they are electronically controlled, add-on units are available for TIG and MIG welding which increase the flexibility. Electrodes are now available in hermetically sealed containers. These vacuum packs obviate the need for baking the electrodes immediately prior to use. However, if a container has been opened or damaged, it is essential that the electrodes are redried according to the manufacturer's instructions.

Equipment for MMA Welding

Although the manual metal arc (MMA) process has relatively basic equipment requirements, it is important that the welder has a knowledge of operating features and performance to comply with welding procedures for the job and, of course, for safety reasons.

Essential equipment

The main components of the equipment required for welding are:

power source electrode holder and cables welder protection fume extractionTools required include: a wire brush to clean the joint area adjacent to the weld (and the weld itself after slag removal); a chipping hammer to remove slag from the weld deposit; and, when removing slag, a pair of clear lens goggles or a face shield to protect the eyes (lenses should be shatter-proof and noninflammable).

Power source

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The primary function of a welding power source is to provide sufficient power to melt the joint. However with MMA the power source must also provide current for melting the end of the electrode to produce weld metal, and it must have a sufficiently high voltage to maintain the arc. A constant current (drooping) characteristic is used.

MMA electrodes are designed to be operated with alternating current (AC) and direct current (DC) power sources. Although AC electrodes can be used on DC, not all DC electrodes can be used with AC power sources.

As MMA requires a high current (50-300A) but a relatively low voltage (10-50V), high voltage mains supply (240 or 440V) must be reduced by a transformer. To produce DC, the output from the transformer must be further rectified. To reduce the hazard of electrical shock, the power source must function with a maximum no-load voltage, that is, when the external (output) circuit is open (power leads connected and live) but no arc is present. The no-load voltage rating of the power source is as defined in IEC 60974-1 and IEC 60974-6 and must be in accordance with the type of welding environment or hazard of electrical shock. The power source may have an internal or external hazard reducing device to reduce the no-load voltage; the main welding current is delivered as soon as the electrode touches the workpiece. For welding in confined spaces, you should use DC or AC with a low voltage safety device to limit the voltage available at the holder to less than 48V rms.

There are four basic types of power source:

AC transformer DC rectifier AC/DC transformer-rectifier DC generatorAC electrodes are frequently operated with the simple, single phase transformer with current adjusted by means of tappings or sliding core control. DC rectifiers and AC/DC transformer-rectifiers are controlled electronically, for example by thyristors. Modern power sources called inverters use transistors to convert mains AC (50Hz) to a high frequency AC (typically 50 kHz) before transforming down to a voltage suitable for welding and then rectifying to DC. Because high frequency transformers can be relatively small, principal advantages of inverter power sources are undoubtedly their size and weight when the source must be portable.

Electrode holder and cables

The electrode holder clamps the end of the electrode with copper contact shoes built into its head. The shoes are actuated by either a twist grip or spring-loaded mechanism. The clamping mechanism allows for quick release of the stub end. For efficiency the electrode has to be firmly clamped into the holder, otherwise poor electrical contact may cause arc instability through voltage fluctuations. Welding cable connecting the holder to the power source is mechanically crimped or soldered. Electrode holders should conform to IEC 60974-11.

It is essential that good electrical connections are maintained between electrode, holder and cable. With poor connections, resistance heating and, in severe cases, minor arcing with the torch body will cause the holder to overheat. Two cables are connected to the output of the power source, the welding lead goes to the electrode holder and the current return lead is clamped to the workpiece. The latter is often wrongly referred to as the earth lead. A separate earth lead may be required to provide protection from faults in the power source. The earth cable should therefore be capable of carrying the maximum output current of the power source. For guidance see DD CLC/TS 62081.

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Cables are covered in a smooth and hard-wearing protective rubberised flexible sheath. This oil and water resistant coating provides electrical insulation at voltages to earth not exceeding 100V DC and AC (rms value). Cable diameter is generally selected on the basis of welding current level. The higher the current and duty cycle, the larger the diameter of the cable to ensure that it does not overheat (see BS 638 Pt 4). If welding is carried out some distance from the power source, it may be necessary to increase cable diameter to reduce voltage drop.

Care of electrodes

The quality of weld relies upon consistent performance of the electrode. The flux coating should not be chipped, cracked or, more importantly, allowed to become damp.

Storage

Electrodes should always be kept in a dry and well-ventilated store. It is good practice to stack packets of electrodes on wooden pallets or racks well clear of the floor. Also, all unused electrodes which are to be returned should be stored so they are not exposed to damp conditions to regain moisture. Good storage conditions are 10 degrees C above external air temperature. As the storage conditions are to prevent moisture from condensing on the electrodes, the electrode stores should be dry rather that warm. Under these conditions and in original packaging, electrode storage time is practically unlimited. It should be noted that electrodes are now available in hermetically sealed packs which obviate the need for drying. However, if necessary, any unused electrodes must be redried according to manufacturer's instructions.

Drying of electrodes

Drying is usually carried out following the manufacturer's recommendations and requirements will be determined by the type of electrode.

Cellulosic coatings

As these electrode coatings are designed to operate with a definite amount of moisture in the coating, they are less sensitive to moisture pick-up and do not generally require a drying operation. However, in cases where ambient relative humidity has been very high, drying may be necessary.

Rutile coatings

These can tolerate a limited amount of moisture and coatings may deteriorate if they are overdried. Particular brands may need to be dried before use.

Basic and basic/rutile coatings

Because of the greater need for hydrogen control, moisture pick-up is rapid on exposure to air. These electrodes should be thoroughly dried in a controlled temperature drying oven. Typical drying time is one hour at a temperature of approximately 150 to 300 degrees C but instructions should be adhered to.

After controlled drying, basic and basic/rutile electrodes must be held at a temperature between 100 and 150 degrees C to help protect them from re-absorbing moisture into the coating. These conditions can be obtained by transferring the electrodes from the main drying oven to a holding oven or a heated quiver at the workplace.

Protective clothing

When welding, the welder must be protected from heat and light radiation emitted from the arc, spatter ejected from the weld pool, and from welding fume.

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Hand and head shield

For most operations a hand-held or head shield constructed of lightweight insulating and non-reflecting material is used which conforms to EN175. The shield is fitted with a protective filter glass, sufficiently dark in colour and capable of absorbing the harmful infrared and ultraviolet rays. The filter glasses conform to the strict requirements of EN 169 and are graded according to a shade number which specifies the amount of visible light allowed to pass through - the lower the number, the lighter the filter. The correct shade number for MMA welding must be used according to the welding current level, for example:

Shade 9 - up to 40A Shade 10 - 40 to 80A Shade 11 - 80 to 175A Shade 12 - 175 to 300A Shade 13 - 300 to 500ANote: The current ranges are different for other processes.

Clothing

For protection against sparks, hot spatter, slag and burns, leather gloves should be worn. Various types of leather gloves are available, such as short or elbow length, full fingered or part mitten. Other protection such as leather aprons may be required depending on the application.

Fume extraction

When welding within a welding shop, ventilation must dispose harmlessly of the welding fume. Particular attention should be paid to ventilation when welding in a confined space such as inside a boiler, tank or compartment of a ship.

Fume removal should be by some form of mechanical ventilation which will produce a current of fresh air in the immediate area. Direction of the air movement should be from the welder's face towards the work. This is best achieved by localised exhaust ventilation using a suitably designed hood near to the welding area.

Oxy-fuel welding

Job Knowledge

Process features

Oxyacetylene welding, commonly referred to as gas welding, is a process which relies on combustion of oxygen and acetylene. When mixed together in correct proportions within a hand-held torch or blowpipe, a relatively hot flame is produced with a temperature of about 3,200 deg.C. The chemical action of the oxyacetylene flame can be adjusted by changing the ratio of the volume of oxygen to acetylene.

Three distinct flame settings are used, neutral, oxidising and carburising. 

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Neutral flame

Oxidising flame

Carburising flame

Welding is generally carried out using the neutral flame setting which has equal quantities of oxygen and acetylene. The oxidising flame is obtained by increasing just the oxygen flow rate while the carburising flame is achieved by increasing acetylene flow in relation to oxygen flow. Because steel melts at a temperature above 1,500 deg.C, the mixture of oxygen and acetylene is used as it is the only gas combination with enough heat to weld steel. However, other gases such as propane, hydrogen and coal gas can be used for joining lower melting point non-ferrous metals, and for brazing and silver soldering.

Equipment

Oxyacetylene equipment is portable and easy to use. It comprises oxygen and acetylene gases stored under pressure in steel cylinders. The cylinders are fitted with regulators and flexible hoses which lead to the blowpipe. Specially designed safety devices such as flame traps are fitted between the hoses and the cylinder regulators. The flame trap prevents flames generated by a 'flashback' from reaching the cylinders; principal causes of flashbacks are the failure to purge the hoses and overheating of the blowpipe nozzle.

When welding, the operator must wear protective clothing and tinted coloured goggles. As the flame is less intense than an arc and very little UV is emitted, general-purpose tinted goggles provide sufficient protection.

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Operating characteristics

The action of the oxyacetylene flame on the surface of the material to be welded can be adjusted to produce a soft, harsh or violent reaction by varying the gas flows. There are of course practical limits as to the type of flame which can be used for welding. A harsh forceful flame will cause the molten weld pool to be blown away, while too soft a flame will not be stable near the point of application. The blowpipe is therefore designed to accommodate different sizes of 'swan neck copper nozzle which allows the correct intensity of flame to be used. The relationship between material thickness, blowpipe nozzle size and welding speed, is shown in the chart. When carrying out fusion welding the addition of filler metal in the form of a rod can be made when required. The principal techniques employed in oxyacetylene welding are leftward, rightward and all-positional rightward. The former is used almost exclusively and is ideally suited for welding butt, fillet and lap joints in sheet thicknesses up to approximately 5mm. The rightward technique finds application on plate thicknesses above 5mm for welding in the flat and horizontal-vertical position. The all-positional rightward method is a modification of the rightward technique and is ideally suited for welding steel plate and in particular pipework where positional welding, (vertical and overhead) has to be carried out. The rightward and all- positional rightward techniques enable the welder to obtain a uniform penetration bead with added control over the molten weldpool and weld metal. Moreover, the welder has a clear view of the weldpool and can work in complete freedom of movement. These techniques are very highly skilled and are less frequently used than the conventional leftward technique.

Equipment for Oxyacetylene Welding

Job Knowledge 

Essential equipment components

Torch

The basic oxyacetylene torch comprises:

torch body (or handle) two separate gas tubes (through the handle connected to the hoses) separate control valves mixer chamber flame tube welding tipNB The cutting torch requires two oxygen supplies to the nozzle, one mixed with fuel gas for preheating and a separate oxygen flow for cutting.

Hoses

Hoses between the torch and the gas regulators should be colour-coded; in the UK: red for acetylene, and blue for oxygen. Fittings on the oxygen hose have right-hand threads; while those on the acetylene hose have left-hand threads.

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Gas regulators

The primary function of a gas regulator is to control gas pressure. It reduces the high pressure of the bottle-stored gas to the working pressure of the torch, and this will be maintained during welding.

The regulator has two separate gauges: a high pressure gauge for gas in the cylinder and a low pressure gauge for pressure of gas fed to the torch. The amount of gas remaining in the cylinder can be judged from the high pressure gauge. The regulator, which has a pressure adjusting screw, is used to control gas flow rate to the torch by setting the outlet gas pressure. Note Acetylene is supplied in cylinders under a pressure of about 15 bar but welding is carried out with torch gas pressures typically up to 2 bar.

Flame traps

Flame traps (also called flashback arresters) must be fitted into both oxygen and acetylene gas lines to prevent a flashback flame from reaching the regulators. Non-return spring-loaded valves can be fitted in the hoses to detect/stop reverse gas flow. Thus, the valves can be used to prevent conditions leading to flashback, but should always be used in conjunction with flashback arresters.

A flashback is where the flame burns in the torch body, accompanied by a whistling sound. It will occur when flame speed exceeds gas flow rate and the flame can pass back through the mixing chamber into the hoses. Most likely causes are: incorrect gas pressures giving too low a gas velocity, hose leaks, loose connections, or welder techniques which disturb gas flow.

Identification of gas cylinders

Gas cylinders are colour-coded. In the UK, an oxygen cylinder is black with a white/grey shoulder; and an acetylene cylinder is maroon. Cylinders should also carry a label that gives details of the type of gas.

Oxygen and acetylene are stored in cylinders at high pressure. Oxygen pressure can be as high as 300 bar. Acetylene, which is dissolved in acetone contained in a porous material, is stored at a much lower pressure, approximately 15bar.

It is vitally important to ensure that the regulator fitted to the oxygen cylinder is rated to at least the same pressure as the cylinder. Some oxygen regulators are only rated at 215 bar and should not be used on a 300 bar cylinder. Flammable gases such as acetylene (and propane) have left hand threads on the cylinder and regulator; the oxygen regulator and cylinder have a conventional right hand thread. On no account should oil or grease be allowed to come into contact with oxygen equipment.

Typical gas pressures and flow rates for C-Mn steel:

Steel thickness (mm)

Nozzle size

Acetylene OxygenPressure

(bar)Consumption

(l/min)Pressure

(bar)Consumption

(l/min)0.90 1 0.14 0.50 0.14 0.501.20 2 0.14 0.90 0.14 0.902.00 3 0.14 1.40 0.14 1.402.60 5 0.14 2.40 0.14 2.403.20 7 0.14 3.30 0.14 3.304.00 10 0.21 4.70 0.21 4.705.00 13 0.28 6.00 0.28 6.006.50 18 0.28 8.50 0.28 8.50

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8.20 25 0.42 12.00 0.42 12.0010.00 35 0.63 17.00 0.63 17.0013.00 45 0.63 22.00 0.63 22.0025.00 90 0.63 42.00 0.63 42.00Selection of correct nozzles

Welding torches are generally rated according to thickness of material to be welded. They range from light duty (for sheet steel up to 2mm in thickness) to heavy duty (for steel plate greater than 25mm in thickness). Each torch can be fitted with a range of nozzles with a bore diameter selected according to material thickness. Gas pressures are set to give correct flow rate for nozzle bore diameter. Proportions of oxygen and acetylene in the mixture can be adjusted to give a neutral, oxidising or carburising flame. (See the description of oxyacetylene processes) Welding is normally carried out using a neutral flame with equal quantities of oxygen and acetylene.

Equipment safety checks

Before commencing welding it is wise to inspect the condition and operation of all equipment. As well as normal equipment and workplace safety checks, there are specific procedures for oxyacetylene. Operators should verify that:

flashback arresters are present in each gas line hoses are the correct colour, with no sign of wear, as short as possible and not taped together regulators are the correct type for the gas a bottle key is in each bottle (unless the bottle has an adjusting screw)It is recommended that oxyacetylene equipment is checked at least annually - regulators should be taken out of service after five years. Flashback arresters should be checked regularly according to manufacturer's instructions and, with specific designs, it may be necessary to replace if flashback has occurred.

MIG / MAG welding

Job Knowledge

Metal inert gas (MIG) welding was first patented in the USA in 1949 for welding aluminium. The arc and weld pool formed using a bare wire electrode was protected by helium gas, readily available at that time. From about 1952, the process became popular in the UK for welding aluminium using argon as the shielding gas, and for carbon steels using CO2. CO2 and argon-CO2 mixtures are known as metal active gas (MAG) processes. MIG is an attractive alternative to MMA, offering high deposition rates and high productivity.

Process characteristics

MIG/MAG welding is a versatile technique suitable for both thin sheet and thick section components. An arc is struck between the end of a wire electrode and the workpiece, melting both of them to form a weld

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pool. The wire serves as both heat source (via the arc at the wire tip) and filler metal for the joint. The wire is fed through a copper contact tube (contact tip) which conducts welding current into the wire. The weld pool is protected from the surrounding atmosphere by a shielding gas fed through a nozzle surrounding the wire. Shielding gas selection depends on the material being welded and the application. The wire is fed from a reel by a motor drive, and the welder moves the welding torch along the joint line. Wires may be solid (simple drawn wires), or cored (composites formed from a metal sheath with a powdered flux or metal filling). Consumables are generally competitively priced compared with those for other processes. The process offers high productivity, as the wire is continuously fed. 

Manual MIG/MAG welding is often referred as a semi-automatic process, as the wire feed rate and arc length are controlled by the power source, but the travel speed and wire position are under manual control. The process can also be mechanised when all the process parameters are not directly controlled by a welder, but might  still require manual adjustment during welding. When no manual intervention is needed during welding, the process can be referred to as automatic.

The process usually operates with the wire positively charged and connected to a power source delivering a constant voltage. Selection of wire diameter (usually between 0.6 and 1.6mm) and wire feed speed determine the welding current, as the burn-off rate of the wire will form an equilibrium with the feed speed.

Metal transfer mode

The manner, or mode, in which the metal transfers from the electrode to the weld pool largely determines the operating features of the process. There are three principal metal transfer modes:

Short circuiting/ Dip Droplet / spray PulsedShort-circuiting and pulsed metal transfer are used for low current operation while spray metal transfer is only used with high welding currents. In short-circuiting or 'dip' transfer, the molten metal forming on the tip of the wire is transferred by the wire dipping into the weld pool. This is achieved by setting a low voltage; for a 1.2mm diameter wire, arc voltage varies from about 17V (100A) to 22V (200A). Care in setting the voltage and the inductance in relation to the wire feed speed is essential to minimise spatter. Inductance is used to control the surge in current which occurs when the wire dips into the weld pool.

For droplet or spray transfer, a much higher voltage is necessary to ensure that the wire does not make contact i.e. short-circuit, with the weld pool; for a 1.2mm diameter wire, the arc voltage varies from approximately 27V (250A) to 35V (400A). The molten metal at the tip of the wire transfers to the weld pool in the form of a spray of small droplets (about the diameter of the wire and smaller). However, there is a minimum current level, threshold, below which droplets are not forcibly projected across the arc. If an open arc technique is attempted much below the threshold current level, the low arc forces would be insufficient

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to prevent large droplets forming at the tip of the wire. These droplets would transfer erratically across the arc under normal gravitational forces. The pulsed mode was developed as a means of stabilising the open arc at low current levels i.e. below the threshold level, to avoid short-circuiting and spatter. Metal transfer is achieved by applying pulses of current, each pulse having sufficient force to detach a droplet. Synergic pulsed MIG refers to a special type of controller which enables the power source to be tuned (pulse parameters) for the wire composition and diameter, and the pulse frequency to be set according to the wire feed speed.

Shielding gas

In addition to general shielding of the arc and the weld pool, the shielding gas performs a number of important functions:

forms the arc plasma stabilises the arc roots on the material surface ensures smooth transfer of molten droplets from the wire to the weld poolThus, the shielding gas will have a substantial effect on the stability of the arc and metal transfer and the behaviour of the weld pool, in particular, its penetration. General purpose shielding gases for MIG welding are mixtures of argon, oxygen and CO2, and special gas mixtures may contain helium. The gases which are normally used for the various materials are:

steels:o CO2

o argon +2 to 5% oxygeno argon +5 to 25% CO2

non-ferrous (e.g. Aluminium, copper or nickel alloys):o argono argon / helium

Argon based gases, compared with CO2 , are generally more tolerant to parameter settings and generate lower spatter levels with the dip transfer mode. However, there is a greater risk of lack of fusion defects because these gases are colder. As CO2 cannot be used in the open arc (pulsed or spray transfer) modes due to high back-plasma forces, argon based gases containing oxygen or CO2 are normally employed.

Applications

MIG/MAG is widely used in most industry sectors and accounts for more than 50% of all weld metal deposited. Compared to MMA, MIG/MAG has the advantage in terms of flexibility, deposition rates and suitability for mechanisation. However, it should be noted that while MIG/MAG is ideal for 'squirting' metal, a high degree of manipulative skill is demanded of the welder.

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Equipment for MIG Welding

The MIG process is a versatile welding technique which is suitable for both thin sheet and thick section components. It is capable of high productivity but the quality of welds can be called into question. To achieve satisfactory welds, welders must have a good knowledge of equipment requirements and should also recognise fully the importance of setting up and maintaining component parts correctly.

Essential equipment

In MIG the arc is formed between the end of a small diameter wire electrode fed from a spool, and the workpiece. Main equipment components are:

power source wire feed system conduit gunThe arc and weldpool are protected from the atmosphere by a gas shield. This enables bare wire to be used without a flux coating (required by MMA). However, the absence of flux to 'mop up' surface oxide places greater demand on the welder to ensure that the joint area is cleaned immediately before welding. This can be done using either a wire brush for relatively clean parts, or a hand grinder to remove rust and scale. The other essential piece of equipment is a wire cutter to trim the end of the electrode wire.

Power source

MIG is operated usually with a DC power source. The source is termed a flat, or constant voltage, characteristic power source, which refers to the voltage/welding current relationship. In MIG, welding current is determined by wire feed speed, and arc length is determined by power source voltage level (open circuit voltage). Wire burn-off rate is automatically adjusted for any slight variation in the gun to workpiece distance, wire feed speed, or current pick-up in the contact tip. For example, if the arc momentarily shortens, arc voltage will decrease and welding current will be momentarily increased to burn back the wire and maintain pre-set arc length. The reverse will occur to counteract a momentary lengthening of the arc.

There is a wide range of power sources available, mode of metal transfer can be:

dip spray pulsedA low welding current is used for thin-section material, or welding in the vertical position. The molten metal is transferred to the workpiece by the wire dipping into the weldpool. As welding parameters will vary from around 100A \17V to 200A \ 22V (for a 1.2mm diameter wire), power sources normally have a current

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rating of up to 350A. Circuit inductance is used to control the surge in current when the wire dips into the weldpool (this is the main cause of spatter). Modern electronic power sources automatically set the inductance to give a smooth arc and metal transfer.

In spray metal transfer, metal transfers as a spray of fine droplets without the wire touching the weldpool. The welding current level needed to maintain the non short-circuiting arc must be above a minimum threshold level; the arc voltage is higher to ensure that the wire tip does not touch the weldpool. Typical welding parameters for a 1.2mm diameter wire are within 250A \ 28V to 400A \ 35V. For high deposition rates the power source must have a much higher current capacity: up to 500A.

The pulsed mode provides a means of achieving a spray type metal transfer at current levels below threshold level. High current pulses between about 25 and 200Hz are used to detach droplets as an alternative to dip transfer. As control of the arc and metal transfer requires careful setting of pulse and background parameters, a more sophisticated power source is required. Synergic pulsed MIG power sources, which are advanced transistor-controlled power sources, are preprogrammed so that the correct pulse parameters are delivered automatically as the welder varies wire feed speed.

Welding current and arc voltage ranges for selected wire diameters operating with dip and spray metal transfer:

Wire diameter (mm)

Dip transfer Spray transferCurrent

(A) Voltage (V) Current (A) Voltage (V)

0.6 30 - 80 15 - 180.8 45 - 180 16 - 21 150 - 250 25 - 331.0 70 - 180 17 - 22 230 - 300 26 - 351.2 100 - 200 17 - 22 250 - 400 27 - 351.6 120 - 200 18 - 22 250 - 500 30 - 40Wire feed system

The performance of the wire feed system can be crucial to the stability and reproducibility of MIG welding. As the system must be capable of feeding the wire smoothly, attention should be paid to the feed rolls and liners. There are three types of feeding systems:

pinch rolls push-pull spool on gunThe conventional wire feeding system normally has a set of rolls where one is grooved and the other has a flat surface. Roll pressure must not be too high otherwise the wire will deform and cause poor current pick up in the contact tip. With copper coated wires, too high a roll pressure or use of knurled rolls increases the risk of flaking of the coating (resulting in copper build up in the contact tip). For feeding soft wires such as aluminium dual-drive systems should be used to avoid deforming the soft wire.

Small diameter aluminium wires, 1mm and smaller, are more reliably fed using a push-pull system. Here, a second set of rolls is located in the welding gun - this greatly assists in drawing the wire through the conduit. The disadvantage of this system is increased size of gun. Small wires can also be fed using a small spool mounted directly on the gun. The disadvantages with this are increased size, awkwardness of the gun, and higher wire cost.

Conduit

The conduit can measure up to 5m in length, and to facilitate feeding, should be kept as short and straight as possible. (For longer lengths of conduit, an intermediate push-pull system can be inserted). It has an internal

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liner made either of spirally-wound steel for hard wires (steel, stainless steel, titanium, nickel) or PTFE for soft wires (aluminium, copper).

Gun

In addition to directing the wire to the joint, the welding gun fulfils two important functions - it transfers the welding current to the wire and provides the gas for shielding the arc and weldpool.

There are two types of welding guns: 'air' cooled and water cooled. The 'air' cooled guns rely on the shielding gas passing through the body to cool the nozzle and have a limited current-carrying capacity. These are suited to light duty work. Although 'air' cooled guns are available with current ratings up to 500A, water cooled guns are preferred for high current levels, especially at high duty cycles.

Welding current is transferred to the wire through the contact tip whose bore is slightly greater than the wire diameter. The contact tip bore diameter for a 1.2mm diameter wire is between 1.4 and 1.5mm. As too large a bore diameter affects current pick up, tips must be inspected regularly and changed as soon as excessive wear is noted. Copper alloy (chromium and zirconium additions) contact tips, harder than pure copper, have a longer life, especially when using spray and pulsed modes.

Gas flow rate is set according to nozzle diameter and gun to workpiece distance, but is typically between 10 and 30 l/min. The nozzle must be cleaned regularly to prevent excessive spatter build-up which creates porosity. Anti-spatter spray can be particularly effective in automatic and robotic welding to limit the amount of spatter adhering to the nozzle.

Protective equipment

Recommended shade number of filter for MIG/MAG welding:

Shade numberWelding current A

MIG Heavy metalMIG Light metal MAG8 up to 70      - up to 709 70 - 125 up to 125 70 - 10010 125 - 175 125 – 175 100 - 15011 175 - 250 175 – 225 150 - 22512 250 - 350 225 – 300 225 - 40013 350 - 450 300 – 400 400 - 60014 450 - 500 400 – 500 over 60015 over 500 over 500      -

MIG/MAG – developments in low heat input transfer modes

Since its initial development in the late 1940s, the MIG/MAG process has seen several iterations in equipment development; which have included control of the way in which metal is transferred from the wire to the weld pool. This mode determines the operating features of the process. Historically, it is considered that there are three principal metal transfer modes:

Short circuiting/ dip Droplet / spray PulsedDip transfer combines low current/heat input and a small wire diameter with repeated short-circuiting between the wire and the weld pool (1), making the process suitable for joining thin sheet and/or positional

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welding, where precise control of the weld pool is required. Recent developments in equipment, associated with advances in inverter technology and electronic control have resulted in greater refinements to the process, including improvements in the control and stability of short circuiting or dip transfer.

Dip transfer

Also known as short circuit metal transfer, this provides the lowest heat input of all transfer modes. The metal transfer occurs when the wire is in contact with the weld pool during the short circuit (2). The current delivered by the power source heats the wire until it begins to melt, during which time the electromagnetic field surrounding the wire increases in strength and creates a force (the pinch effect) which separates the molten part from the rest of the electrode (Figure 1). After this, the wire melts into the weld pool and the cycle begins again.

Figure 1: Dip transfer before and during the separation of a molten drop of metal from the wire

The steps in the sequence for conventional dip transfer are illustrated in Figure 2. When the electrode touches the weld pool (A) a short circuit is created, the arc extinguishes, the voltage decreases and the current increases into the short circuit. This causes the drop to be released (B). The arc reignites when the contact between the wire and the weld pool is broken (C). The cycle then starts again with the arc reignition (D) and re-melting and wire contact (E) .The frequency of the cycle is typically 50 to 150Hz.

Figure 2: Current and voltage fluctuations during the dip-transfer mode of MIG/MAG welding

Dip transfer, combines low current/heat input and a small wire diameter with repeated short-circuiting between the wire and the weld pool (2), making the process suitable for joining thin sheet and/or positional

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welding, where precise control of the weld pool is required. The main disadvantage associated with conventional dip transfer is the high spatter level associated with the fluctuations of the current and voltage cycle. In order to improve the quality of the weld and the efficiency of the process it is necessary to increase the control on the current cycle applied. This improvement has come about with the development of the most recent power sources.

Equipment developments

There have been a large number of developments by welding equipment manufacturers to improve process stability and reduce spatter, giving a large choice of potential systems that fabricators can adopt.

Spatter is associated with molten metal being squeezed by a pinch force during the current rise–droplet detachment phases of the process cycle. The rate of current rise is critical, in order to balance maintaining a molten wire for metal transfer against excess current/pinch force, and subsequent spatter. To this end, different equipment manufacturers have come up with a number of solutions; all of which relate to improving the control and stability of the current profile, with an aim to reducing spatter and improving heat input control. To do this, all of the systems rely on development of digitally controlled power sources, and a more precise control of the waveform. Such systems include the Fronius Cold Metal Transfer (CMT), EWM ColdArc®, Lincoln Electric Surface Tension Transfer (STT), Miller Regulated Metal Deposition (RMD™), Kemppi FastROOT, Jetline Controlled Short Circuit (CSC™) MIG, Daihen Corp. Controlled Bridge Transfer (CBT), Merkle ColdMIG  and ESAB QSet™ processes.

The majority of these systems control the process by electronic regulation within the power supply. Whilst each has its own particular current profile characteristics, all rely on a rapid reduction of the welding current immediately prior to arc re-ignition (4-8) which can take place in a more controlled manner compared with conventional dip transfer (see Figure 3). As a result, manufacturers claim significant reductions in spatter, with a 5-30% reduction in heat input, permitting joining of material thicknesses as low as 0.3mm (4,5), and high gap bridging ability (up to 4.8mm, (9)).

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Figure 3 Typical waveforms produced by different power sources. Clockwise from top-left: EWM ColdArc, Miller RMD, Daihen CBT, Merkle ColdMIG and Lincoln STT

The only major difference between most of the systems relates to whether they are regulated by software or hardware. The exception is the Fronius CMT system, which integrates control of the motion of the wire into the welding process control to support droplet formation and detachment (10,11); see Figure 4. Wire feed towards the workpiece (A) is reversed when short circuiting occurs (B), at which point the wire is retracted (C). Metal transfer supported by surface tension in the melt means that the current can be maintained at a very low level; with reduced heat input and spatter (12). After opening of the short circuit, the wire speed is changed back to feed into the weld pool (D). In this case, the process cycle is random, with the oscillation frequency varying with time; but typically around 70 Hz (5). In addition to the power source, there is the additional requirement of a special wire feed unit and torch. Whilst it has been reported that the system can be adopted for manual welding (5), the majority of applications are mechanised or automated, partly as a consequence of the need to manipulate a larger, heavier torch (4).

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Figure 4 Principal phases of wire feed control in the CMT process

Summary

When asked to state their needs for further developments for MIG/MAG processes, fabricators expressed the following as critical: reduction of spatter (on dip transfer), improved arc stability, lower non-fusion defects, increased tolerance to wide gaps and positional welding ability. As such, these have featured heavily as the primary advantages of by most recent short-arc process variants. There are now several MIG/MAG short-arc welding manufacturer specific variants available and well known by the users in industry which offer users some if not all of these capabilities.

Submerged-arc Welding

The first patent on the submerged-arc welding (SAW) process was taken out in 1935 and covered an electric arc beneath a bed of granulated flux. Developed by the E O Paton Electric Welding Institute, Russia, during the Second World War, SAW's most famous application was on the T34 tank.

Process features

Similar to MIG welding, SAW involves formation of an arc between a continuously-fed bare wire electrode and the workpiece. The process uses a flux to generate protective gases and slag, and to add alloying elements to the weld pool. A shielding gas is not required. Prior to welding, a thin layer of flux powder is placed on the workpiece surface. The arc moves along the joint line and as it does so, excess flux is recycled via a hopper. Remaining fused slag layers can be easily removed after welding. As the arc is completely covered by the flux layer, heat loss is extremely low. This produces a thermal efficiency as high as 60% (compared with 25% for manual metal arc). There is no visible arc light, welding is spatter-free and there is no need for fume extraction.

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Operating characteristics

SAW is usually operated as a fully-mechanised or automatic process, but it can be semi-automatic. Welding parameters: current, arc voltage and travel speed all affect bead shape, depth of penetration and chemical composition of the deposited weld metal. Because the operator cannot see the weld pool, greater reliance must be placed on parameter settings.

Process variants

According to material thickness, joint type and size of component, varying the following can increase deposition rate and improve bead shape.

Wire

SAW is normally operated with a single wire on either AC or DC current. Common variants are:

twin wire multiple wire (tandem or triple) single wire with hot or cold wire addition metal powder addition tubular wireAll contribute to improved productivity through a marked increase in weld metal deposition rates and/or travel speeds.

A narrow gap process variant is also established, which utilises a two or three bead per layer deposition technique.

What is narrow gap welding?

Narrow gap welding (also called narrow groove welding) was developed to weld thick sections more economically. This welding procedure uses joint preparations with small, included angles, typically in the range 2-20°, which require less weld metal and less welding time to fill. Narrow gap techniques have been applied when welding using submerged arc welding (SAW), gas shielded metal arc welding (MIG/MAG, GMAW) and tungsten inert gas welding (TIG, GTAW) processes. However, narrow gap welding does require specialised equipment, because of the limited accessibility to the root of the preparation.

The advantages of the narrow gap technique are:

The process offers better economy for welding of thick materials (generally over 50mm thick) because of reduced consumable requirements and shorter welding times.

There is low angular distortion because the joint preparation is almost parallel-sided.The disadvantages are:

The weld is more prone to defects for certain welding processes - especially lack of sidewall fusion.

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It is difficult to remove any defects when detected, because of poor joint accessibility. Expensive J-preparations must be machined onto the parent material, unless a backing bar is

permitted. This will affect the economics of the process.The risk of lack of sidewall fusion can be reduced in narrow gap welding by several methods:

1. Using two electrodes in tandem with each electrode oriented so that a weld bead is directed towards each sidewall (applicable to SAW and MIG/MAG processes)

2. Using an electrode that has been bent into a wave form (for MIG/MAG welding). This should make the arc move from side to side across the joint

3. Using two electrodes that are twisted around each other to oscillate the arc (applicable to MIG/MAG welding)

4. Using an angled contact tip, which automatically aims the electrode at one sidewall and then the other (applicable to MIG/MAG welding)

5. Arc oscillation

6. Use of seam tracking to ensure alignment of the arc with the sidewall

Flux

Fluxes used in SAW are granular fusible minerals containing oxides of manganese, silicon, titanium, aluminium, calcium, zirconium, magnesium and other compounds such as calcium fluoride. The flux is specially formulated to be compatible with a given electrode wire type so that the combination of flux and wire yields desired mechanical properties. All fluxes react with the weld pool to produce the weld metal chemical composition and mechanical properties. It is common practice to refer to fluxes as 'active' if they add manganese and silicon to the weld, the amount of manganese and silicon added is influenced by the arc voltage and the welding current level. The the main types of flux for SAW are:

Bonded fluxes - produced by drying the ingredients, then bonding them with a low melting point compound such as a sodium silicate. Most bonded fluxes contain metallic deoxidisers which help to prevent weldporosity. These fluxes are effective over rust and mill scale.

Fused fluxes - produced by mixing the ingredients, then melting them in an electric furnace to form a chemically homogeneous product, cooled and ground to the required particle size. Smooth stable arcs, with welding currents up to 2000A and consistent weld metal properties, are the main attraction of these fluxes.

Applications

SAW is ideally suited for longitudinal and circumferential butt and fillet welds. However, because of high fluidity of the weld pool, molten slag and loose flux layer, welding is generally carried out on butt joints in the flat position and fillet joints in both the flat and horizontal-vertical positions. For circumferential joints, the workpiece is rotated under a fixed welding head with welding taking place in the flat position. Depending on material thickness, either single-pass, two-pass or multipass weld procedures can be carried out. There is virtually no restriction on the material thickness, provided a suitable joint preparation is adopted. Most commonly welded materials are carbon-manganese steels, low alloy steels and stainless steels, although the process is capable of welding some non-ferrous materials with judicious choice of electrode filler wire and flux combinations.

Equipment for Submerged-arc Welding

Job Knowledge

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The submerged-arc welding (SAW) process is similar to MIG where the arc is formed between a continuously-fed wire electrode and the workpiece, and the weld is formed by the arc melting the workpiece and the wire. However, in SAW a shielding gas is not required as the layer of flux generates the gases and slag to protect the weld pool and hot weld metal from contamination. Flux plays an additional role in adding alloying elements to the weld pool.

Essential equipment

Essential equipment components for SAW are:

power source SAW head flux handling protective equipmentAs SAW is a high current welding process, the equipment is designed to produce high deposition rates.

Power source

SAW can be operated using either a DC or an AC power source. DC is supplied by a transformer-rectifier and AC is supplied by a transformer. Current for a single wire ranges from as low as 200A (1.6mm diameter wire) to as high as 1000A (6.0mm diameter wire). In practice, most welding is carried out on thick plate where a single wire (4.0mm diameter) is normally used over a more limited range of 600 to 900A, with a twin wire system operating between 800 and 1200A.

In DC operation, the electrode is normally connected to the positive terminal. Electrode negative (DCEN) polarity can be used to increase deposition rate but depth of penetration is reduced by between 20 and 25%. For this reason, DCEN is used for surfacing applications where parent metal dilution is important. The DC power source has a 'constant voltage' output characteristic which produces a self-regulating arc. For a given diameter of wire, welding current is controlled by wire feed speed and arc length is determined by voltage setting.

AC power sources usually have a constant-current output characteristic and are therefore not self-regulating. The arc with this type of power source is controlled by sensing the arc voltage and using the signal to control wire feed speed. In practice, for a given welding current level, arc length is determined by wire burnoff rate, i.e. the balance between the welding current setting and wire feed speed which is under feedback control.

Square wave AC square wave power sources have a constant voltage output current characteristic. Advantages are easier arc ignition and constant wire feed speed control.

Welding gun

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SAW can be carried out using both manual and mechanised techniques. Mechanised welding, which can exploit the potential for extremely high deposition rates, accounts for the majority of applications.

Manual welding

For manual welding, the welding gun is similar to a MIG gun, with the flux which is fed concentrically around the electrode, replacing the shielding gas. Flux is fed by air pressure through the handle of the gun or from a small hopper mounted on the gun. The equipment is relatively portable and, as the operator guides the gun along the joint, little manipulative skill is required. However, because the operator has limited control over the welding operation (apart from adjusting travel speed to maintain the bead profile) it is best used for short runs and simple filling operations.

Mechanised welding - single wire

As SAW is often used for welding large components, the gun, wire feeder and flux delivery feed can be mounted on a rail, tractor or boom manipulator. Single wire welding is mostly practised using DCEP even though AC will produce a higher deposition rate for the same welding current. AC is used to overcome problems with arc blow, caused by residual magnetism in the workpiece, jigging or welding machine.

Wire stickout, or electrode extension - the distance the wire protrudes from the end of the contact tip - is an important control parameter in SAW. As the current flowing between the contact tip and the arc will preheat the wire, wire burnoff rate will increase with increase in wire stickout. For example, the deposition rate for a 4mm diameter wire at a welding current of 700A can be increased from approximately 9 kg/hr at the normal 32mm stickout, to 14 kg/hr at a stickout length of 178mm. In practice, because of the reduction in penetration and greater risk of arc wander, a long stickout is normally only used in cladding and surfacing applications where there is greater emphasis on deposition rate and control of penetration, rather than accurate positioning of the wire.

For most applications, electrode stickout is set so that the contact tube is slightly proud of the flux layer. The depth of flux is normally just sufficient to cover the arc whose light can be seen through the flux.

Recommended and maximum stickout lengths:

Wire diameter mm Current range A Wire stickout

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Normal mm Maximum mm

0.8 100 to 200 12 -

1.2 150 to 300 20 -

1.6 200 to 500 20 -

2.0 250 to 600 25 63

3.2 350 to 800 30 76

4.0 400 to 900 32 128

4.75 450 to 1000 35 165

Mechanised welding - twin wire

Tandem arc connections

SAW can be operated with more than one wire. Although up to five wires are used for high deposition rates, e.g. in pipe mills, the most common multi-wire systems have two wires in a tandem arrangement. The leading wire is run on DCEP to produce deep penetration. The trailing wire is operated on AC which spreads the weld pool, which is ideal for filling the joint. AC also minimises: interaction between the arcs, and the risk of lack of fusion defects and porosity through the deflection of the arcs (arc blow). The wires are normally spaced 20mm apart so that the second wire feeds into the rear of the weld pool.

Gun angle

In manual welding, the gun is operated with a trailing angle, i.e. with the gun at an angle of 45 degrees (backwards) from the vertical. In single wire mechanised welding operations, the gun is perpendicular to the workpiece. However, in twin wire operations the leading gun is normal to the workpiece, with the trailing gun angled slightly forwards between an angle of 60 and 80 degrees. This reduces disturbance of the weld pool and produces a smooth weld bead profile.

Flux handling

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Flux should be stored in unopened packages under dry conditions. Open packages should be stored in a humidity-controlled store. While flux from a newly-opened package is ready for immediate use, flux which has been opened and held in a store should first be dried according to manufacturer's instructions. In small welding systems, flux is usually held in a small hopper above the welding gun. It is fed automatically (by gravity or mechanised feed) ahead of the arc. In larger installations the flux is stored in large hoppers and is fed with compressed air. Unused flux is collected using a vacuum hose and returned to the hopper.

Note: Care must be taken in recycling unused flux, particularly regarding the removal of slag and metal dust particles. The presence of slag will change the composition of the flux which, together with the wire, determines the composition of the weld metal. The presence of fine particles can cause blockages in the feeding system.

Protective equipment

Unlike other arc welding processes, SAW is a clean process which produces minimum fume and spatter when welding steels. (Some noxious emissions can be produced when welding special materials.) For normal applications, general workshop extraction should be adequate.

Protective equipment such as a head shield and a leather apron are not necessary. Normal protective equipment (goggles, heavy gloves and protective shoes) are required for ancillary operations such as slag removal by chipping or grinding. Special precautions should be taken when handling flux - a dust respirator and gloves are needed when loading the storage hoppers.

TIG Welding

Tungsten inert gas (TIG) welding became an overnight success in the 1940s for joining magnesium and aluminium. Using an inert gas shield instead of a slag to protect the weldpool, the process was a highly attractive replacement for gas and manual metal arc welding. TIG has played a major role in the acceptance of aluminium for high quality welding and structural applications.

Process characteristics

In the TIG process the arc is formed between a pointed tungsten electrode and the workpiece in an inert atmosphere of argon or helium. The small intense arc provided by the pointed electrode is ideal for high quality and precision welding. Because the electrode is not consumed during welding, the welder does not have to balance the heat input from the arc as the metal is deposited from the melting electrode. When filler metal is required, it must be added separately to the weldpool.

Power source

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TIG must be operated with a drooping, constant current power source - either DC or AC. A constant current power source is essential to avoid excessively high currents being drawn when the electrode is short-circuited on to the workpiece surface. This could happen either deliberately during arc starting or inadvertently during welding. If, as in MIG welding, a flat characteristic power source is used, any contact with the workpiece surface would damage the electrode tip or fuse the electrode to the workpiece surface. In DC, because arc heat is distributed approximately one-third at the cathode (negative) and two-thirds at the anode (positive), the electrode is always negative polarity to prevent overheating and melting. However, the alternative power source connection of DC electrode positive polarity has the advantage in that when the cathode is on the workpiece, the surface is cleaned of oxide contamination. For this reason, AC is used when welding materials with a tenacious surface oxide film, such as aluminium.

Arc starting

The welding arc can be started by scratching the surface, forming a short-circuit. It is only when the short-circuit is broken that the main welding current will flow. However, there is a risk that the electrode may stick to the surface and cause a tungsten inclusion in the weld. This risk can be minimised using the 'lift arc' technique where the short-circuit is formed at a very low current level. The most common way of starting the TIG arc is to use HF (High Frequency). HF consists of high voltage sparks of several thousand volts which last for a few microseconds. The HF sparks will cause the electrode - workpiece gap to break down or ionise. Once an electron/ion cloud is formed, current can flow from the power source.

Note: As HF generates abnormally high electromagnetic emission (EM), welders should be aware that its use can cause interference especially in electronic equipment. As EM emission can be airborne, like radio waves, or transmittedalong power cables, care must be taken to avoid interference with control systems and instruments in the vicinity of welding.

HF is also important in stabilising the AC arc; in AC, electrode polarity is reversed at a frequency of about 50 times per second, causing the arc to be extinguished at each polarity change. To ensure that the arc is reignited at each reversal of polarity, HF sparks are generated across the electrode/workpiece gap to coincide with the beginning of each half-cycle.

Electrodes

Electrodes for DC welding are normally pure tungsten with 1 to 4% thoria to improve arc ignition. Alternative additives are lanthanum oxide and cerium oxide which are claimed to give superior performance (arc starting and lower electrode consumption). It is important to select the correct electrode diameter and tip angle for the level of welding current. As a rule, the lower the current the smaller the electrode diameter and tip angle. In AC welding, as the electrode will be operating at a much higher temperature, tungsten with a zirconia addition is used to reduce electrode erosion. It should be noted that because of the large amount of heat generated at the electrode, it is difficult to maintain a pointed tip and the end of the electrode assumes a spherical or 'ball' profile.

Shielding gas

Shielding gas is selected according to the material being welded. The following guidelines may help: Argon - the most commonly-used shielding gas which can be used for welding a wide range of

materials including steels, stainless steel, aluminium and titanium. Argon + 2 to 5% H2 - the addition of hydrogen to argon will make the gas slightly reducing,

assisting the production of cleaner-looking welds without surface oxidation. As the arc is hotter and more constricted, itpermits higher welding speeds. Disadvantages include risk of hydrogen cracking in carbon steels and weld metal porosity in aluminium alloys.

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Helium and helium/argon mixtures - adding helium to argon will raise the temperature of the arc. This promotes higher welding speeds and deeper weld penetration. Disadvantages of using helium or a helium/argon mixtureis the high cost of gas and difficulty in starting the arc.

Applications

TIG is applied in all industrial sectors but is especially suitable for high quality welding. In manual welding, the relatively small arc is ideal for thin sheet material or controlled penetration (in the root run of pipe welds). Because deposition rate can be quite low (using a separate filler rod) MMA or MIG may be preferable for thicker material and for fill passes in thick-wall pipe welds.

TIG is also widely applied in mechanised systems either autogenously or with filler wire. However, several 'off the shelf' systems are available for orbital welding of pipes, used in the manufacture of chemical plant or boilers. The systems require no manipulative skill, but the operator must be well trained. Because the welder has less control over arc and weldpool behaviour, careful attention must be paid to edge preparation (machined rather than hand-prepared), joint fit-up and control of welding parameters.

Equipment for TIG Welding

Job Knowledge

Using an inert gas shield instead of a slag to protect the weldpool, this technology is a highly attractive alternative to gas and manual metal arc welding and has played a major role in the acceptance of high quality welding in critical applications.

Essential equipment

In TIG welding, the arc is formed between the end of a small diameter tungsten electrode and the workpiece. The main equipment components are:

power source torch backing system protective equipment Power source

The power source for TIG welding can be either DC or AC, but in both the output is termed a drooping, or constant current characteristic; the arc voltage/welding current relationship delivers a constant current for a given power source setting. 

In TIG welding, the arc length is dependent on how consistently the welder can hold the torch above the workpiece. Arc length is directly proportional to arc voltage, so a longer arc has a higher voltage and if the arc is shortened, the voltage will decrease. Variation of arc length by 3 or 4mm can easily vary the voltage by 5V. By design, the TIG power source has a limited range of current and a reduced variation on changing

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voltage. With such a power source, the variation of current over a variation of 5V might be as little as 10A, giving almost imperceptible changes to the weld pool, making control much easier for the welder.

The arc is usually started by High Frequency (HF) sparks which ionise the gap between the electrode and the workpiece. HF generates airborne and line transmitted interference, so care must be taken to avoid interference with control systems and instruments near welding equipment. When welding is carried out in sensitive areas, a non-HF technique, touch starting or 'lift arc', can be used. The electrode can be short circuited to the workpiece, but the current will only flow when the electrode is lifted off the surface. There is, therefore, little risk of the electrode fusing to the workpiece surface and forming tungsten inclusions in the weld metal. For high quality applications, using HF is preferred.

DC power source

DC power produces a concentrated arc with most of the heat in the workpiece, so this power source is generally used for welding. However, the arc with its cathode roots on the electrode (DC electrode negative polarity), results in little cleaning of the workpiece surface. Care must be taken to clean the surface prior to welding and to ensure that there is an efficient gas shield.

Transistor and inverter power sources are being used increasingly for TIG welding. The advantages are:

the smaller size makes them easily transported arc ignition is easier special operating features (e.g. current pulsing) are readily included the output can be pre-programmed for mechanised operationsThe greater stability of these power sources allows very low currents to be used particularly for micro-TIG welding and largely replaced the plasma process for micro-welding operations.

AC power source

For materials such as aluminium, which has a tenacious oxide film on the surface, AC power must be employed. By switching between positive and negative polarity, the periods of electrode positive will remove the oxide and clean the surface.

The figure shows current and voltage waveforms for (sine wave) AC TIG welding.

Disadvantages of conventional sine wave AC compared with DC are:

More diffuse arc HF is required to reignite the arc at each current reversal

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Excessive heating of the electrode makes it impossible to maintain a tapered point and the end becomes balled

Square wave AC, or switched DC, power sources are particularly attractive for welding aluminium. 

By switching between polarities, arc reignition is made easier so that the HF can be reduced or eliminated. The ability to imbalance the waveform to vary the proportion of positive to negative polarity is important by determining the relative amount of heat generated in the workpiece and the electrode.

To weld the root run, the power source is operated with the greater amount of positive polarity to put the maximum heat into the workpiece. 

For filler runs a greater proportion of negative polarity should be used to minimise heating of the electrode. By using 90% negative polarity, it is possible to maintain a pointed electrode. A balanced position (50% electrode positive and negative polarities) is preferable for welding heavily oxidised aluminium.

Torch

There is a wide range of torch designs for welding, depending on the application. Designs which have the on/off switch and current control in the handle are often preferred to foot controls. Specialised torches are available for mechanised applications, e.g. orbital and bore welding of pipes.

Electrode

The electrode tip is usually ground to an angle of 60 to 90 degrees for manual welding, regardless of the electrode diameter. For mechanised applications as the tip angle determines the shape of the arc and influences the penetration profile of the weld pool, attention must be paid to consistency in grinding the tip and checking its condition between welds.

For AC current, the electrode often pure tungsten. The tip normally adopts a spherical profile due to the heat generated in the electrode during the electrode positive half cycle.

Gas shielding

A gas lens should be fitted within the torch nozzle, to ensure laminar gas flow. This will improve gas protection for sensitive welding operations like welding vertical, corner and edge joints and on curved surfaces. There is also a wide range of nozzles available, ensuring different gas coverage. The nozzle's selection depends mainly on the electrode diameter and on the accessibility, defined by the assembly to be welded.

Backing system

When welding high integrity components, a shielding gas is used to protect the underside of the weld pool and weld bead from oxidation. To reduce the amount of gas consumed, a localised gas shroud for sheet, dams or plugs for tubular components is used. As little as 5% air can result in a poor weld bead profile and may reduce corrosion resistance in materials like stainless steel. With gas backing systems in pipe welding, pre-weld purge time depends on the diameter and length of the pipe. The flow rate and purge time are set to ensure at least five volume changes before welding.

Stick on tapes and ceramic backing bars are also used to protect and support the weld bead. In manual stainless steel welding, a flux-cored wire instead of a solid wire can be used in the root run. This protects the underbead from oxidation without the need for gas backing.

Inserts

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A pre-placed insert can be used to improve the uniformity of the root penetration. Its main use is to prevent suck-back in an autogenous weld, especially in the overhead position. The use of an insert does not make welding any easier and skill is still required to avoid problems of incomplete root fusion and uneven root penetration.

Protective equipment

A slightly darker glass should be used in the head or hand shield than that used for MMA welding.Recommended shade number of filter for TIG welding:

Shade number Welding current A9 less than 2010 20 to 4011 40 to 10012 100 to 17513 175 to 25014 250 to 400

Plasma Arc Welding

Job Knowledge

Process characteristics

Plasma welding is very similar to TIG as the arc is formed between a pointed tungsten electrode and the workpiece. However, by positioning the electrode within the body of the torch, the plasma arc can be separated from the shielding gas envelope. Plasma is then forced through a fine-bore copper nozzle which constricts the arc. Three operating modes can be produced by varying bore diameter and plasma gas flow rate:

Microplasma: 0.1 to 15A.The microplasma arc can be operated at very low welding currents. The columnar arc is stable even when arc length is varied up to 20mm.

Medium current: 15 to 200A.At higher currents, from 15 to 200A, the process characteristics of the plasma arc are similar to the TIG arc, but because the plasma is constricted, the arc is stiffer. Although the plasma gas flow rate can be

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increased to improve weld pool penetration, there is a risk of air and shielding gas entrainment through excessive turbulence in the gas shield.

Keyhole plasma: over 100A.By increasing welding current and plasma gas flow, a very powerful plasma beam is created which can achieve full penetration in a material, as in laser or electron beam welding. During welding, the hole progressively cuts through the metal with the molten weld pool flowing behind to form the weld bead under surface tension forces. This process can be used to weld thicker material (up to 10mm of stainless steel) in a single pass.

Power source

The plasma arc is normally operated with a DC, drooping characteristic power source. Because its unique operating features are derived from the special torch arrangement and separate plasma and shielding gas flows, a plasma control console can be added on to a conventional TIG power source. Purpose-built plasma systems are also available. The plasma arc is not readily stabilised with sine wave AC. Arc reignition is difficult when there is a long electrode to workpiece distance and the plasma is constricted, Moreover, excessive heating of the electrode during the positive half-cycle causes balling of the tip which can disturb arc stability.

Special-purpose switched DC power sources are available. By imbalancing the waveform to reduce the duration of electrode positive polarity, the electrode is kept sufficiently cool to maintain a pointed tip and achieve arc stability.

Arc starting

Although the arc is initiated using HF, it is first formed between the electrode and plasma nozzle. This 'pilot' arc is held within the body of the torch until required for welding then it is transferred to the workpiece. The pilot arc system ensures reliable arc starting and, as the pilot arc is maintained between welds, it obviates the need for HF which may cause electrical interference.

Electrode

The electrode used for the plasma process is tungsten-2%thoria and the plasma nozzle is copper. The electrode tip diameter is not as critical as for TIG and should be maintained at around 30-60 degrees. The plasma nozzle bore diameter is critical and too small a bore diameter for the current level and plasma gas flow rate will lead to excessive nozzle erosion or even melting. It is prudent to use the largest bore diameter for the operating current level.

Note: too large a bore diameter, may give problems with arc stability and maintaining a keyhole.

Plasma and shielding gases

The normal combination of gases is argon for the plasma gas, with argon plus 2 to 5% hydrogen for the shielding gas. Helium can be used for plasma gas but because it is hotter this reduces the current rating of the nozzle. Helium's lower mass can also make the keyhole mode more difficult.

Applications

Microplasma welding

Microplasma was traditionally used for welding thin sheets (down to 0.1 mm thickness), and wire and mesh sections. The needle-like stiff arc minimises arc wander and distortion. Although the equivalent TIG arc is more diffuse, the newer transistorised (TIG) power sources can produce a very stable arc at low current levels.

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Medium current welding

When used in the melt mode this is an alternative to conventional TIG. The advantages are deeper penetration (from higher plasma gas flow), and greater tolerance to surface contamination including coatings (the electrode is within the body of the torch). The major disadvantage lies in the bulkiness of the torch, making manual welding more difficult. In mechanised welding, greater attention must be paid to maintenance of the torch to ensure consistent performance.

Keyhole welding

This has several advantages which can be exploited: deep penetration and high welding speeds. Compared with the TIG arc, it can penetrate plate thicknesses up to l0mm, but when welding using a single pass technique, it is more usual to limit the thickness to 6mm. The normal methods is to use the keyhole mode with filler to ensure smooth weld bead profile (with no undercut). For thicknesses up to 15mm, a vee joint preparation is used with a 6mm root face. A two-pass technique is employed and here, the first pass is autogenous with the second pass being made in melt mode with filler wire addition.

As the welding parameters, plasma gas flow rate and filler wire addition (into the keyhole) must be carefully balanced to maintain the keyhole and weld pool stability, this technique is only suitable for mechanised welding. Although it can be used for positional welding, usually with current pulsing, it is normally applied in high speed welding of thicker sheet material (over 3 mm) in the flat position. When pipe welding, the slope-out of current and plasma gas flow must be carefully controlled to close the keyhole without leaving a hole.

Equipment for Plasma Welding

Plasma welding derives its unique operating characteristics from the torch design. As in TIG welding, the arc is formed between the end of a small diameter tungsten electrode and the workpiece. However, in the plasma torch, the electrode is positioned behind a fine bore copper nozzle. By forcing the arc to pass through the nozzle, the characteristic columnar jet, or plasma, is formed.

As described in Job Knowledge for Welders, No 7, three different operating modes can be produced by the choice of the nozzle bore diameter, current level and plasma gas flow rate:

Microplasma (0.1 to 15A) is equivalent to microTIG but the columnar arc allows the welder to operate with a much longer arc length. The arc is stable at low welding current levels producing a 'pencil-like' beam which is suitable for welding very thin section material.

Medium current plasma (15 to 200A) similar to conventional TIG, is also used for precision welding operations and when a high level of weld quality is demanded.

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Keyhole plasma (over 100A) produced by increasing the current level and the plasma gas flow. It generates a very powerful arc plasma, similar to a laser beam. During welding, the plasma arc slices through the metal producing a keyhole, with the molten weld pool flowing around the keyhole to form the weld. Deep penetration and high welding speeds can be achieved with this operating mode.

As the plasma arc is generated by the special torch arrangement and system controller, the equipment can be obtained as an add-on unit to conventional TIG equipment to provide additional pilot arc and separate plasma and shielding gases. Alternatively, purpose-built plasma equipment is available. Despite similarities in plasma and TIG equipment, there are several important differences in the following components: 

power source torch backing system protective equipmentPower source

The power source for plasma welding is almost exclusively DC and, as in TIG, the drooping, or constant current, output characteristic will deliver essentially constant current for a given power source setting. The power source is ideal for mechanised welding as it maintains the current setting even when arc length varies and, in manual welding, it can accommodate the natural variations of the welder.

The plasma process is normally operated with electrode negative polarity to minimise heat produced in the electrode (approximately 1/3rd of the heat generated by the arc is produced at the cathode with 2/3rds at the anode). Special torches are available, however, for operating with electrode positive polarity which rely on efficient cooling to prevent melting of the electrode. The positive electrode torch is used for welding aluminium which requires the cathode to be on the material to remove the oxide film.

AC is not normally used in the plasma process because it is difficult to stabilise the AC arc. Problems in reigniting the arc are associated with constriction by the nozzle, the long electrode to workpiece distance and balling of the electrode caused by the alternate periods of electrode positive polarity. The square wave AC (inverter, switched DC) power source, with an efficiently cooled torch, makes the use of the AC plasma process easier; rapid current switching promotes arc reignition and, by operating with very short periods of electrode positive polarity, electrode heating is reduced so a pointed electrode can be maintained.

The plasma system has a unique arc starting system in which HF is only used to ignite a pilot arc held within the body of the torch. The pilot arc formed between the electrode and copper nozzle is automatically transferred to the workpiece when it is required for welding. This starting system is very reliable and eliminates the risk of electrical interference through HF.

Torch

The torch for the plasma process is considerably more complex than the TIG torch and attention must be paid, not only to initial set up, but also to inspection and maintenance during production.

Nozzle

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In the conventional torch arrangement, the electrode is positioned behind the water cooled copper nozzle. As the power of the plasma arc is determined by the degree of nozzle constriction, consideration must be given to the choice of bore diameter in relation to the current level and plasma gas flow rate. For a 'soft' plasma, normally used for micro and medium current operating modes, a relatively large diameter bore is recommended to minimise nozzle erosion.

In high current keyhole plasma mode, the nozzle bore diameter, plasma gas flow rate and current level are selected to produce a highly constricted arc which has sufficient power to cut through the material. The plasma gas flow rate is crucial in generating the deeply penetrating plasma arc and in preventing nozzle erosion; too low a gas flow rate for the bore diameter and current level will result in double arcing in the torch and the nozzle melting.

The suggested starting point for setting the plasma gas flow rate and the current level for a range of the bore diameters and the various operating modes is given.

Electrode

The electrode is tungsten with an addition of between 2 and 5% thoria to aid arc initiation. Normally, the electrode tip is ground to an angle of 15 degrees for microplasma welding. The tip angle increases with current level and for high current, keyhole plasma welding, an angle of 60 degrees to 90 degrees is recommended. For high current levels, the tip is also blunted to approximately 1mm diameter. The tip angle is not usually critical for manual welding. However, for mechanised applications, the condition of the tip and the nozzle will determine the shape of the arc and penetration profile of the weld pool penetration, so particular attention must be paid to grinding the tip. It is also necessary to check periodically the condition of the tip and nozzle and, for critical components, it is recommended the torch condition is checked between welds. 

Electrode set-back

To ensure consistency, it is important to maintain a constant electrode position behind the nozzle; guidance on electrode set-back and a special tool is provided by the torch manufacturer. The maximum current rating of each nozzle has been established for the maximum electrode set-back position and the maximum plasma gas flow rate. Lower plasma gas flow rates can be used to soften the plasma arc with the maximum current rating of the nozzle providing electrode set-back distance is reduced.

Plasma and shielding gas

The usual gas combination is argon for the plasma gas and argon-2 to 8% H2 for the shielding gas. Irrespective of the material being welded, using argon for the plasma gas produces the lowest rate of electrode and nozzle erosion. Argon - H2 gas mixture for shielding produces a slightly reducing atmosphere and cleaner welds. Helium gives a hotter arc; however, its use for the plasma gas reduces the current

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carrying capacity of the nozzle and makes formation of the keyhole more difficult. Helium - argon mixtures, e.g. 75% helium - 25% argon, are used as the shielding gas for materials such as copper.

Plasma gas flow rate must be set accurately as it controls the penetration of the weld pool but the shielding gas flow rate is not critical.

Backing system

The normal TIG range of backing bar designs or shielding gas techniques can be employed when using micro and medium current techniques. When applying the keyhole mode a grooved backing bar must be used, with or without gas shielding or total shielding of the underside of the joint. Because the efflux plasma normally extends about 10mm below the back face of the joint, the groove must be deep enough to avoid disturbance of the arc jet; if the efflux plasma hits the backing bar, arc instability will disturb the weld pool, causing porosity.

Protective equipment

Protective equipment for plasma welding is as described for TIG in Job Knowledge for Welders No 17. Regarding protection from arc light, a similar Shade number to TIG at the same welding current level should be used in head or hand shield. The glass will be slightly darker than that used for MMA welding at the same current level.  

Recommended shade number of filter for plasma welding:

Shade NunberWelding Current, A

Micro Plasma Plasma5 0.5 to 16 1 to 2.57 2.5 to 58 5 to 109 10 to 1510 15 to 3011 30 to 60 less than 15012 60 to 125 150 to 25013 125 to 225 above 25014 225 to 450

Thermal Gouging

Job Knowledge

Thermal gouging is an essential part of welding fabrication. Used for rapid removal of unwanted metal, the material is locally heated and molten metal ejected - usually by blowing it away. Normal oxyfuel gas or arc processes can be used to produce rapid melting and metal removal. However, to produce a groove of specific dimensions, particularly regarding depth and width, the welder must exercise careful control of the gouging operation. If this does not happen, an erratic and badly-serrated groove will result. 

Thermal processes, operations and metals which may be gouged or otherwise shaped:

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Thermal process

Process operationsMetals

Primary Secondary

Oxyfuel gas flame Gouging

Grooving Washing Chamfering

Low carbon steels, carbon manganese steels (structural), pressure vessel steels (carbon not over 0.35%), low alloy steels (less than 5%Cr) cast iron (if preheated to 400-450 deg.C)

Manual metal arc Gouging Grooving 

ChamferingLow carbon steels carbon manganese steels (structural), pressure vessel steels, low alloy steels, stainless steels, cast iron, nickel-based alloys

Air carbon arc Gouging Grooving 

Chamfering

Low carbon steels carbon manganese steels (structural), pressure vessel steels, low and high alloy steels, cast iron, nickel-based alloys, copper and copper alloys, copper/nickel alloys, aluminium

Plasma arc GougingChamferingGrooving Washing

Aluminium, stainless steels

Note: All processes are capable of cutting/severing operations. Preheat may or may not be required on some metals prior to gougingSafety

It should be emphasised that because gouging relies on molten metal being forcibly ejected, often over quite large distances, the welder must take appropriate precautions to protect himself, other workers and his equipment. Sensible precautions include protective clothing for the welder, shielding inside a specially-enclosed booth or screens, adequate fume extraction, and removal of all combustible material from the immediate area.

Industrial applications

Thermal gouging was developed primarily for removal of metal from the reverse side of welded joints, removal of tack welds, temporary welds, and weld imperfections. Figure 1 illustrates the value of typical back-gouging applications carried out on arc welded joints., while Fig. 2 shows imperfection removal in preparation for weld repair. 

Fig.1 Typical back-gouging applications carried out on arc welded joints

Fig. 2 Imperfection removal in preparation for weld repair

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The gouging process has proved to be so successful that it is used for a wide spectrum of applications in engineering industries:

repair and maintenance of structures - bridges, earth-moving equipment, mining machinery, railway rolling stock, ships, offshore rigs, piping and storage tanks

removal of cracks and imperfections - blow holes and sand traps in both ferrous and non-ferrous forgings and castings

preparation of plate edges for welding

removal of surplus metal - riser pads and fins on castings, excess weld bead profiles, temporary backing strips, rivet washing and shaping operations, demolition of welded and unwelded structures - site work

Thermal gouging is also suitable for efficient removal of temporary welded attachments such as brackets, strongbacks, lifting lugs and redundant tack welds, during various stages of fabrication and construction work.

Gouging processes

Gouging operations can be carried out using the following thermal processes: oxyfuel gas flame manual metal arc air carbon arc plasma arc

Oxy-fuel Gouging

Job Knowledge Oxy-fuel or flame gouging offers fabricators a quick and efficient method of removing metal. It can

be at least four times quicker than cold chipping operations. The process is particularly attractive because of its low noise, ease of handling, and ability to be used in all positions.

Process description Flame gouging is a variant of conventional oxyfuel gas welding. Oxygen and a fuel gas are used to

produce a high temperature flame for melting the steel. When gouging, the steel is locally heated to a temperature above the 'ignition' temperature (typically 900deg.C) and a jet of oxygen is used to melt the metal - a chemical reaction between pure oxygen and hot metal. This jet is also used to blow away molten metal and slag. It should be noted that compared with oxyfuel cutting, slag is not blown through the material, but remains on the top surface of the workpiece.

The gouging nozzle is designed to supply a relatively large volume of oxygen through the gouging jet. This can be as much as 300 litre/min through a 6mm orifice nozzle. In oxyacetylene gouging, equal quantities of oxygen and acetylene are used to set a near-neutral preheating flame. The oxygen jet flow rate determines the depth and width of the gouge. Typical operating parameters (gas pressures and flow rates) for achieving a range of gouge sizes (depth and width) can be seen in the Table.

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Typical operating data for manual oxyacetylene flame gouging

Nozzle orifice dia.

(mm)

Gouge dimensions Gas pressure Gas consumption Travel

speed (mm/min)Width

(mm)Depth (mm)

Acetylene (Bar)

Oxygen (Bar)

Acetylene (Litre/min)

Preheat (Litre/min)

Oxygen (Litre/min)

3 6-8 3-9 0.48 4.2 15 22 62 6005 8-10 6-12 0.48 5.2 29 31 158 10006.5 10-13 10-13 0.55 5.5 36 43 276 1200

When the preheating flame and oxygen jet are correctly set, the gouge has a uniform profile and its surfaces are smooth with a dull blue colour.

Operating techniques The depth of the gouge is determined principally by the speed and angle of the torch. To cut a deep

groove the angle of the torch is stepped up (this increases the impingement angle of the oxygen jet) and gouging speed is reduced. To produce a shallow groove, the torch is less steeply angled, see above, and speed is increased. Wide grooves can be produced by weaving the torch. The contour of the groove is dependent upon the size of the nozzle and the operating parameters. If the cutting oxygen pressure is too low, gouging progresses with a washing action, leaving smooth ripples in the bottom of the groove. If the cutting oxygen pressure is too high, the cut advances ahead of the molten pool - this will disrupt the gouging operation especially when making shallow grooves.

There are four basic flame gouging techniques which are used in the following types of application. Progressive gouging  This technique is used to produce uniform grooves. Gouging is conducted in either a continuous or

progressive manner. Applications include removal of an unfused root area on the reverse side of a welded joint, part-shaping a steel forging, complete removal of a weld deposit and preparing plate edges for welding.

Spot gouging Spot gouging produces a deep narrow U-shaped groove over a relatively short length. The process is

ideally suited to removal of localised areas such as isolated weld imperfections. Experienced operators are able to observe any imperfections during gouging. These appear as dark or light spots/streaks within the molten pool (reaction zone).

Back-step gouging Once the material has reached ignition temperature, the oxygen stream is introduced and the torch

moved in a backward movement for a distance of 15-20mm. The oxygen is shut off and the torch moved forward a distance of 25-30mm before restarting the gouging operation. This technique is favoured for removal of local imperfections which may be deeply embedded in the base plate.

Deep gouging It is sometimes necessary to produce a long deep gouge. Such operations are completed using the

deep gouging technique, which is basically a combination of progressive and spot gouging.

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Manual Metal Arc Gouging Job Knowledge The main advantage of manual metal arc (MMA) gouging is that the same power source can be used

for welding, gouging, or cutting, simply by changing the type of electrode. Process description

As in conventional MMA welding, the arc is formed between the tip of the electrode and the

workpiece. MMA gouging differs because it requires special purpose electrodes with thick flux coatings to generate a strong arc force and gas stream. Unlike MMA welding where a stable weld pool must be maintained, this process forces the molten metal away from the arc zone to leave a clean cut surface.

The gouging process is characterised by the large amount of gas which is generated to eject the molten metal. However, because the arc/gas stream is not as powerful as a gas or a separate air jet, the surface of the gouge is not as smooth as an oxyfuel gouge or air carbon arc gouge.

Electrode According to the size of gouge specified, there is a wide range of electrode diameters available to

choose from. These grooving electrodes are also not just restricted to steels, and the same electrode composition may be used for gouging stainless steel and non-ferrous alloys.

Power source MMA gouging can be carried out using conventional DC and AC power sources. In DC gouging,

electrode polarity is normally negative but electrode manufacturers may well recommend electrode polarity for their brand of electrodes and for gouging specific materials. When using an AC power source, a minimum of 70V open circuit (OCV) is required to stabilise the arc.

Although most MMA welding power sources can be used for gouging, the current rating and OCV must be capable of accommodating current surges and longer arc lengths.

Guidance on gouging parameters can be found below:

Typical operating data for MMA gouging

Electrode diameter (mm) Current (A)

Gouging dimensionsGouging speed (mm/min)

Depth (mm) Width (mm)3.2 210 2 6 12004.0 300 3 8 10004.8 350 4 10 800

Operational characteristics The arc is struck with an electrode which is held at a normal angle to the workpiece (15 degrees

backwards from the vertical plane in line with proposed direction of gouging). Once the arc is established, the electrode is immediately inclined in one smooth and continuous movement to an angle of around 15-20 degrees to the plate surface. With the arc pointing in the direction of travel, the electrode is pushed forward slightly to melt the metal. It should then be pulled back to allow the gas jet to displace the molten metal and slag. This forward and backward motion is repeated as the electrode is guided along the line to complete the gouge.

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To produce a consistent depth and width of gouge, a uniform rate of travel must be maintained, together with the angle of electrode: 10-20 degrees. If the electrode angle becomes too steep, in excess of about 20 degrees, the amount of slag and molten metal will increase. This is a result of the arc penetrating too deeply. Digging the electrode into the metal causes problems in controlling the gouging operation and will produce a rough surface profile. For gouging in positions other than vertical, the electrode is always pushed forward. With vertical surfaces, the electrode is directed and pushed vertically downwards.

Application MMA gouging is used for localised gouging operations, removal of defects for example, and where

it is more convenient to switch from a welding electrode to a gouging electrode rather than use specialised equipment. Compared with alternative gouging processes, metal removal rates are low and the quality of the gouged surface is inferior. 

When correctly applied, MMA gouging can produce relatively clean gouged surfaces. For general applications, welding can be carried out without the need to dress by grinding. However when gouging stainless steel, a thin layer of higher carbon content material will be produced - this should be removed by grinding.

Plasma Arc Gouging Job Knowledge The use of the plasma arc as a gouging tool dates back to the 1960s when the process was developed

for welding. Compared with the alternative oxyfuel and MMA gouging techniques, plasma arc has a needle-like jet which can produce a very precise groove, suitable for application on almost all ferrous and non- ferrous materials.

Process description

Plasma arc gouging is a variant of the plasma arc process. The arc is formed between a refractory

(usually tungsten) electrode and the workpiece. Intense plasma is achieved by constricting the arc using a fine bore copper nozzle. By locating the electrode behind the nozzle, the plasma-forming gas can be separated from the general gas supply used to cool the torch/assist the plasma gas to blow away molten metal (dross) from the groove.

The temperature and force of the constricted plasma arc is determined by the current level and plasma gas flow rate. Thus, the plasma can be varied to produce a hot gas stream or a high power, deeply penetrating jet. This ability to control quite precisely the size and shape of a groove is very useful for removing unwanted defects from a workpiece surface.

Whilst gouging, normal precautions should be taken to protect the operator and other workers in the immediate area from the effects of intense arc light and hot metal spray. Unlike the oxyfuel and MMA processes, the plasma arc's high velocity jet will propel fume and hot metal dross some considerable distance from the operator. When using a deeply penetrating arc, noise protection is an essential requirement.

Equipment

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The power source for sustaining this gouging arc must have a high open circuit voltage, usually well in excess of 100V. The torch is connected to the negative polarity of the power source and the workpiece must be connected to the positive. The plasma torch is the same as the one used for cutting; it will be either gas or water cooled and have the facility for single and dual gas operation.

Electrodes are normally tungsten for argon and argon-based gases. However, when using air as the plasma gas, special purpose, for example hafnium tipped copper, electrodes must be used to withstand the more aggressive, oxidising arc.

Plasma and cooling gases Plasma gas can be argon, helium, argon - H2 , nitrogen or air. Argon - 35% H2 is normally

recommended as a general- purpose plasma gas for cutting most materials. Alternative plasma gases are argon and helium. Argon, a colder gas, will reduce metal removal rates. Helium, which generates a hot but less intense arc than argon - H2 , can produce a wider and shallower groove. Nitrogen and air are also used as plasma gases, especially for gouging C-Mn steels. Although gas costs will be substantially reduced, the groove surface profile will be inferior to that which can be achieved with argon - H2 gas. Air is not recommended for gouging aluminium as this requires an inert or reducing gas. Argon, nitrogen or air are all used as cooling gases. Use of argon will normally produce the best quality of gouge, but nitrogen or air will reduce operating costs.

Operating techniques Gouging is effected by moving the torch forward at a steady controlled rate. It is carried out in a

progressive manner to remove metal over a distance of 200 to 250mm. The jet can then be repositioned, either to deepen or widen the groove, or to continue gouging for a further 200 to 250mm. Principal process parameters are current level, gas flow rate, and speed of gouging. These settings determine groove size and metal removal rate. In a typical gouging operation on C-Mn steel, metal is removed at about 100 kg/hr at a speed of 0.5 m/min, and groove size will be around 12mm wide and 5mm deep.

The torch stand-off and its angle to the surface of the workpiece have a major influence on speed of travel, groove profile and quality of surface. The torch is normally held at a distance of 20mm from the workpiece and inclined backwards to the direction of gouging at an angle of 40 to 45 degrees. Gouging will remove up to approximately 6mm depth of metal in a single pass.

The torch stand-off should not be reduced to less than 12mm, to avoid spatter build-up on the nozzle from the molten particles ejected from the groove. At standoff distances greater than 25mm, arc/gas forces are reduced and this lessens the depth of penetration of the jet. By reducing the torch angle to the workpiece surface, the plasma jet can be encouraged to 'skate' along the surface of the workpiece; this produces a shallower and wider groove. By increasing the angle of the torch the plasma jet is directed into the workpiece surface, resulting in a deeper and narrower groove.

Air Carbon Arc Gouging

The main difference between this gouging technique and the others is that a separate air jet is used to eject molten metal from the groove.

Process description

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Air carbon arc gouging works as follows. An electric arc is generated between the tip of a carbon electrode and the workpiece. The metal becomes molten and a high velocity air jet streams down the electrode to blow it away, thus leaving a clean groove. The process is simple to apply (using the same equipment as MMA welding), has a high metal removal rate, and gouge profile can be closely controlled. Disadvantages are that the air jet causes the molten metal to be ejected over quite a large distance and, because of high currents (up to 2000A) and high air pressures (80 to 100 psi), it can be very noisy.

Application As air carbon arc gouging does not rely on oxidation it can be applied to a wide range of metals. DC

(electrode positive) is normally preferred for steel and stainless steel but AC is more effective for cast iron, copper and nickel alloys. Typical applications include back gouging, removal of surface and internal defects, removal of excess weld metal and preparation of bevel edges for welding.

Electrode

The electrode is a graphite (carbon) rod which has a copper coating to reduce electrode erosion.

Electrode diameter is selected according to required depth and width of gouge. Cutting can be precisely controlled and molten metal/dross is kept to a minimum.

Power source A DC power supply with electrode positive polarity is most suitable. AC power sources which are

also constant current can be used but with special AC type electrodes. The power source must have a constant current output characteristic. If it does not, inadvertant touching of the electrode to the workpiece will cause a high current surge sufficient to 'explode' the electrode tip. This will disrupt the operation and cause carbon pick-up. As arc voltage can be quite high (up to 50V), open circuit voltage of the power source should be over 60V.

Air supply The gouging torch is normally operated with either a compressed air line or separate bottled gas

supply. Air supply pressure will be up to 100psi from the air line but restricted to about 35psi from a bottled supply. Providing there is sufficient air flow to remove molten metal, there are no advantages in using higher pressure and flow rates.

Carbon pickup Although carbon is picked up by the molten metal, the air stream will remove carbon-rich metal

from the groove to leave only minimal contamination of the side walls. Poor gouging technique or

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insufficient air flow will result in carbon pick-up with the risk of metallurgical problems, e.g high hardness and even cracking.

Typical operating data for air carbon arc gouging:

Electrode diameter (mm)

Current ANote: DC electrode

Gouging dimensions Carbon electrode

consumed (mm/min)Gouging speed

(mm/min)Depth (mm)

Width (mm)

Manual

6.4 275 6-7 9-10 120 6098.0 350 7-8 10-11 114 7119.5 425 9-10 12-13 100 66013.0 550 12-13 18-19 76 508

Automatic

8.0 300-400 2-9 3-8 100 1650-8409.5 500 3-12 3-10 142 1650-63513.0 850 3-15 3-13 82 1830-61016.0 1250 3-19 3-16 63 1830-710

Operation Gouging is commenced by striking the electrode tip on to the workpiece surface to initiate the arc.

Unlike manual metal arc (MMA) welding the electrode tip is not withdrawn to establish arc length. Molten metal directly under the electrode tip (arc) is immediately blown away by the air stream. For effective metal removal, it is important that the air stream is directed at the arc from behind the electrode and sweeps under the tip of the electrode. The width of groove is determined by the diameter of electrode, but depth is dictated by the angle of electrode to the workpiece and rate of travel. Relatively high travel speeds are possible when a low electrode angle is used. This produces a shallow groove: a steep angle results in a deep groove and requires slower travel speed. Note, a steeply angled electrode may give rise to carbon contamination.

Oscillating the electrode in a circular or restricted weave motion during gouging can greatly increase gouging width. This is useful for removal of a weld or plate imperfection that is wider than the electrode itself. It is important, however, that weave width should not exceed four times the diameter of the electrode. The groove surface should be relatively free of oxidised metal and can be considered ready for welding without further preparation. Dressing by grinding the side-walls of the gouge should be carried out if a carbon rich layer has been formed. Also, dressing by grinding or another approved method will be necessary if working on crack-sensitive material such as high strength, low alloy steel.

SteelsWeldability of materials

In arc welding, as the weld metal needs mechanical properties to match the parent metal, the welder must avoid forming defects in the weld. Imperfections are principally caused by:

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poor welder technique; insufficient measures to accommodate the material or welding process; high stress in the component.Techniques to avoid imperfections such as lack of fusion and slag inclusions, which result from poor welder techniques, are relatively well known. However, the welder should be aware that the material itself may be susceptible to formation of imperfections caused by the welding process. In the materials section of the Job Knowledge for Welders, guidelines are given on material weldability and precautions to be taken to avoid defects.

Material types

In terms of weldability, commonly used materials can be divided into the following types:

Steels Stainless steels Aluminium and its alloys Nickel and its alloys Copper and its alloys Titanium and its alloys Cast ironFusion welding processes can be used to weld most alloys of these materials, in a wide range of thickness. When imperfections are formed, they will be located in either the weld metal or the parent material immediately adjacent to the weld, called the heat affected zone (HAZ). As chemical composition of the weld metal determines the risk of imperfections, the choice of filler metal may be crucial not only in achieving adequate mechanical properties and corrosion resistance but also in producing a sound weld. However, HAZ imperfections are caused by the adverse effect of the heat generated during welding and can only be avoided by strict adherence to the welding procedure.

This part of the materials section of Job Knowledge for Welders considers the weldability of carbon-manganese (C-Mn) steels and low alloy steels.

Imperfections in welds

Commonly used steels are considered to be readily welded. However, these materials can be at risk from the following types of imperfection:

porosity ;  solidification cracking ; hydrogen cracking ; reheat cracking .Other fabrication imperfections are lamellar tearing and liquation cracking but using modern steels and consumables, these types of defects are less likely to arise.

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In discussing the main causes of imperfections, guidance is given on procedure and welder techniques for reducing the risk in arc welding.

Porosity

Porosity is formed by entrapment of discrete pockets of gas in the solidifying weld pool. The gas may originate from poor gas shielding, surface contaminants such as rust or grease, or insufficient deoxidants in the parent metal (autogenous weld), electrode or filler wire. A particularly severe form of porosity is 'wormholes', caused by gross surface contamination or welding with damp electrodes.

The presence of manganese and silicon in the parent metal, electrode and filler wire is beneficial as they act as deoxidants combining with entrapped air in the weld pool to form slag. Rimming steels with a high oxygen content, can only be welded satisfactorily with a consumable which adds aluminium to the weld pool.

To obtain sound porosity-free welds, the joint area should be cleaned and degreased before welding. Primer coatings should be removed unless considered suitable for welding by that particular process and procedure. When using gas shielded processes, the material surface demands more rigorous cleaning, such as by degreasing, grinding or machining, followed by final degreasing, and the arc must be protected from draughts.

Solidification cracking

Solidification cracks occur longitudinally as a result of the weld bead having insufficient strength to withstand the contraction stresses within the weld metal. Sulphur, phosphorus, and carbon pick up from the parent metal at high dilution increase the risk of weld metal (solidification) cracking especially in thick section and highly restrained joints. When welding high carbon and sulphur content steels, thin weld beads will be more susceptible to solidification cracking. However, a weld with a large depth to width ratio can also be susceptible. In this case, the centre of the weld, the last part to solidify, will have a high concentration of impurities increasing the risk of cracking.

Solidification cracking is best avoided by careful attention to the choice of consumable, welding parameters and welder technique. To minimise the risk, consumables with low carbon and impurity levels and relatively high manganese and silicon contents are preferred. High current density processes such as submerged-arc and CO 2 , are more likely to induce cracking. The welding parameters must produce an adequate depth to width ratio in butt welds, or throat thickness in fillet welds. High welding speeds also increase the risk as the amount of segregation and weld stresses will increase. The welder should ensure that there is a good joint fit-up so as to avoid bridging wide gaps. Surface contaminants, such as cutting oils, should be removed before welding.

Hydrogen cracking

A characteristic feature of high carbon and low alloy steels is that the HAZ immediately adjacent to the weld hardens on welding with an attendant risk of cold (hydrogen) cracking. Although the risk of cracking is determined by the level of hydrogen produced by the welding process, susceptibility will also depend upon several contributory factors:

material composition (carbon equivalent); section thickness; arc energy (heat) input; degree of restraint.The amount of hydrogen generated is determined by the electrode type and the process. Basic electrodes generate less hydrogen than rutile electrodes (MMA) and the gas shielded processes (MIG and TIG)

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produce only a small amount of hydrogen in the weld pool. Steel composition and cooling rate determines the HAZ hardness. Chemical composition determines material hardenability, and the higher the carbon and alloy content of the material, the greater the HAZ hardness. Section thickness and arc energy influences the cooling rate and hence, the hardness of the HAZ.

For a given situation therefore, material composition, thickness, joint type, electrode composition and arc energy input, HAZ cracking is prevented by heating the material. Using preheat which reduces the cooling rate, promotes escape of hydrogen and reduces HAZ hardness so preventing a crack-sensitive structure being formed; the recommended levels of preheat for various practical situations are detailed in the appropriate standards e.g. BS EN1011-2:2001. As cracking only occurs at temperatures slightly above ambient, maintaining the temperature of the weld area above the recommended level during fabrication is especially important. If the material is allowed to cool too quickly, cracking can occur up to several hours after welding, often termed 'delayed hydrogen cracking'. After welding, therefore, it is beneficial to maintain the heating for a given period (hold time), depending on the steel thickness, to enable the hydrogen to diffuse from the weld area.

When welding C-Mn structural and pressure vessel steels, the measures which are taken to prevent HAZ cracking will also be adequate to avoid hydrogen cracking in the weld metal. However, with increasing alloying of the weld metal e.g. when welding alloyed or quenched and tempered steels, more stringent precautions may be necessary.

The risk of HAZ cracking is reduced by using a low hydrogen process, low hydrogen electrodes and high arc energy, and by reducing the level of restraint. Practical precautions to avoid hydrogen cracking include drying the electrodes and cleaning the joint faces. When using a gas shielded process, a significant amount of hydrogen can be generated from contaminants on the surface of the components and filler wire so preheat and arc energy requirements should be maintained even for tack welds.

Reheat cracking

Reheat or stress relaxation cracking may occur in the HAZ of thick section components, usually of greater than 50mm thickness. The more likely cause of cracking is embrittlement of the HAZ during high temperature service or stress relief heat treatment.

As a coarse grained HAZ is more susceptible to cracking, low arc energy input welding procedures reduce the risk. Although reheat cracking occurs in sensitive materials, avoidance of high stresses during welding and elimination of local points of stress concentration, e.g. by dressing the weld toes, can reduce the risk.

Weldability of steel groups

PD CEN ISO/TR 15608:2005 identifies a number of steels groups which have similar metallurgical and welding characteristics. The main risks in welding these groups are:

Group 1. Low carbon unalloyed steels, no specific processing requirements, specified minimum yield strength R eH ≤ 460N/mm 2 .

For thin section, unalloyed materials, these are normally readily weldable. However, when welding thicker sections with a flux process, there is a risk of HAZ hydrogen cracking, which will need increased hydrogen control of the consumables or the use of preheat.

Group 2. Thermomechanically treated fine grain steels and cast steels with a specified miniumum yield strength R eH > 360N/mm 2 .

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For a given strength level, a thermomechanically processed ( TMCP) steel will have a lower alloy content than a normalised steel, and thus will be more readily weldable with regard to avoidance of HAZ hydrogen cracking and the achievement of maximum hardness limits. However, there is always some degree of softening in the HAZ after welding TMCP steels, and a restriction on the heat input used, so as not to degrade the properties of the joint zone (e.g. ≤2.5kJ/mm limits for 15mm plate).

Group 3. Quenched and tempered steels and precipitation hardened steels (except stainless steels), ReH>360N/mm2

These are weldable, but care must be taken to adhere to established procedures, as these often have high carbon contents, and thus high hardenability, leading to a hard HAZ susceptibility to cracking. As with TMCP steels, there maybe a restriction on heat input or preheat to avoid degradation of the steel properties.

Groups 4, 5 and 6. Chromium-molybdenum and chromium-molybdenum-vanadium creep resisting steels.

These are susceptible to hydrogen cracking, but with appropriate preheat and low hydrogen consumables, with temper bead techniques to minimise cracking, the steels are fairly weldable. Postweld heat treatment is used to improve HAZ toughness in these steels.

Group 7. Ferritic, martensitic or precipitation hardened stainless steels.

When using a filler to produce matching weld metal strength, preheat is needed to avoid HAZ cracking. Postweld heat treatment is essential to restore HAZ toughness.

Group 8. Austenitic stainless steels.

These steels do not generally need preheat, but in order to avoid problems with solidification or liquation cracking upon welding, the consumables should be selected to give weld metal with a low impurity content, or if appropriate, residual ferrite in the weld metal.

Group 9. Nickel alloy steels, Ni≤10%.

These have a similar weldability to Groups 4, 5 & 6.

Group 10. Austenitic ferritic stainless steels (duplex).

In welding these steels, maintaining phase balance in the weld metal and in the HAZ requires careful selection of consumables, the absence of preheat and control of maximum interpass temperature, along with minimum heat input levels, as slow cooling encourages austenite formation in the HAZ.

Group 11. High carbon steels.

These steels will be less weldable owing to their increased carbon content with respect to Group 1. It is likely that care over the choice of consumables and the use of high preheat levels would be needed.

It is important to obtain advice before welding any steels that you do not have experience in.

Stainless steelWeldability of materials

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Stainless steels are chosen because of their enhanced corrosion resistance, high temperature oxidation resistance or their strength. The various types of stainless steel are identified and guidance given on welding processes and techniques which can be employed in fabricating stainless steel components without impairing the corrosion, oxidation and mechanical properties of the material or introducing defects into the weld.

Material types

The unique properties of the stainless steels are derived from the addition of alloying elements, principally chromium and nickel, to steel. Typically, more than 10% chromium is required to produce a stainless iron. The four grades of stainless steel have been classified according to their material properties and welding requirements:

Austenitic Ferritic Martensitic Austenitic-ferritic (duplex)The alloy groups are designated largely according to their microstructure. The first three consist of a single phase but the fourth group contains both ferrite and austenite in the microstructure.

As nickel (plus carbon, manganese and nitrogen) promotes austenite and chromium (plus silicon, molybdenum and niobium) encourages ferrite formation, the structure of welds in commercially available stainless steels can be largely predicted on the basis of their chemical composition. The predicted weld metal structure is shown in the Schaeffler diagram in which austenite and ferrite promoting elements are plotted in terms of the nickel and chromium equivalents.

Because of the different microstructures, the alloy groups have both different welding characteristics and susceptibility to defects.

Austenitic stainless steel

Austenitic stainless steels typically have a composition within the range 16-26% chromium (Cr) and 8-22% nickel (Ni). A commonly used alloy for welded fabrications is Type 304 which contains approximately 18%Cr and 10%Ni. These alloys can be readily welded using any of the arc welding processes (TIG, MIG, MMA and SA). As they are non-hardenable on cooling, they exhibit good toughness and there is no need for pre- or post-weld heat treatment.

Avoiding weld imperfections

Although austenitic stainless steel is readily welded, weld metal and HAZ cracking can occur. Weld metal solidification cracking is more likely in fully austenitic structures which are more crack sensitive than those containing a small amount of ferrite. The beneficial effect of ferrite has been attributed largely to its

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capacity to dissolve harmful impurities which would otherwise form low melting point segregates and interdendritic cracks.

As the presence of 5-10% ferrite in the microstructure is extremely beneficial, the choice of filler material composition is crucial in suppressing the risk of cracking. An indication of the ferrite-austenite balance for different compositions is provided by the Schaeffler diagram. For example, when welding Type 304 stainless steel, a Type 308 filler material which has a slightly different alloy content, is used.

Ferritic stainless steel

Ferritic stainless steels have a Cr content typically within the range 11-28%. Commonly used alloys include the 430 grade, having 16-18% Cr and 407 grade having 10-12% Cr. As these alloys can be considered to be predominantly single phase and non-hardenable, they can be readily fusion welded. However, a coarse grained HAZ will have poor toughness.

Avoiding weld imperfections

The main problem when welding this type of stainless steel is poor HAZ toughness. Excessive grain coarsening can lead to cracking in highly restrained joints and thick section material. When welding thin section material, (less than 6mm) no special precautions are necessary.

In thicker material, it is necessary to employ a low heat input to minimise the width of the grain coarsened zone and an austenitic filler to produce a tougher weld metal. Although preheating will not reduce the grain size, it will reduce the HAZ cooling rate, maintain the weld metal above the ductile-brittle transition temperature and may reduce residual stresses. Preheat temperature should be within the range 50-250 deg. C depending on material composition.

Martensitic stainless steel

The most common martensitic alloys e.g. type 410, have a moderate chromium content, 12-18% Cr, with low Ni but more importantly have a relatively high carbon content. The principal difference compared with welding the austenitic and ferritic grades of stainless steel is the potentially hard HAZ martensitic structure and the matching composition weld metal. The material can be successfully welded providing precautions are taken to avoid cracking in the HAZ, especially in thick section components and highly restrained joints.

Avoiding weld imperfections

High hardness in the HAZ makes this type of stainless steel very prone to hydrogen cracking. The risk of cracking generally increases with the carbon content. Precautions which must be taken to minimise the risk, include:

using low hydrogen process (TIG or MIG) and ensure the flux or flux coated consumable are dried (MMA and SAW) according to the manufacturer's instructions;

preheating to around 200 to 300 deg. C. Actual temperature will depend on welding procedure, chemical composition (especially Cr and C content), section thickness and the amount of hydrogen entering the weld metal;

maintaining the recommended minimum interpass temperature. carrying out post-weld heat treatment, e.g. at 650-750 deg. C. The time and temperature will be

determined by chemical composition.Thin section, low carbon material, typically less than 3mm, can often be welded without preheat, providing that a low hydrogen process is used, the joints have low restraint and attention is paid to cleaning the joint area. Thicker section and higher carbon (> 0.1%) material will probably need preheat and post-weld heat

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treatment. The post-weld heat treatment should be carried out immediately after welding not only to temper (toughen) the structure but also to enable the hydrogen to diffuse away from the weld metal and HAZ.

Duplex stainless steels

Duplex stainless steels have a two phase structure of almost equal proportions of austenite and ferrite. The composition of the most common duplex steels lies within the range 22-26% Cr, 4-7% Ni and 0-3% Mo normally with a small amount of nitrogen (0.1-0.3%) to stabilise the austenite. Modern duplex steels are readily weldable but the procedure, especially maintaining the heat input range, must be strictly followed to obtain the correct weld metalstructure.

Avoiding weld imperfections

Although most welding processes can be used, low heat input welding procedures are usually avoided. Preheat is not normally required and the maximum interpass temperature must be controlled. Choice of filler is important as it is designed to produce a weld metal structure with a ferrite-austenite balance to match the parent metal. To compensate for nitrogen loss, the filler may be overalloyed with nitrogen or the shielding gas itself may contain a small amount of nitrogen.

Welding of austenitic stainless steelThere are a number of different types of steels that may be referred to as 'stainless'; previous articles have considered ferritic and precipitation hardening steels for example. It is therefore advisable to be specific and to refer to the group to which the steel belongs in order to avoid confusion. Although commonly referred to as 'stainless steel', the steels covered in this article should be more correctly referred to as austenitic, 18/8 or chromium-nickel stainless steels.

As with the other types of stainless steels, the austenitic stainless steels are corrosion and oxidation resistant due to the presence of chromium that forms a self-healing protective film on the surface of the steel. They also have very good toughness at extremely low temperatures so are used extensively in cryogenic applications. They can be hardened and their strength increased by cold working but not by heat treatment. They are the most easily weldable of the stainless steel family and can be welded by all welding processes, the main problems being avoidance of hot cracking and the preservation of corrosion resistance.

A convenient and commonly used shorthand identifying the individual alloy within the austenitic stainless steel group is the ASTM system. This uses a three digit number '3XX', the '3' identifying the steel as an austenitic stainless, and with additional letters to identify the composition and certain characteristics of the alloy eg type 304H, type 316L etc; this ASTM method will be used in this article.

Typical compositions of some of the alloys are given in Table 1. The type 304 grade may be regarded as the archetypal austenitic stainless steel from which the other grades are derived and changes in composition away from that of type 304 result in a change in the identification number and are highlighted in red.

Table 1 Typical compositions of some austenitic stainless steel alloys

ASTM No.(type)

Composition wt% Microstructure

C (max) Si (max) Mn (max) Cr Ni Mo Others Austenite - A

Ferrite - F304 0.08 0.75 2.0 18/20 8/11 - - A+2/8%F

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304L 0.035 0.75 2.0 18/20 8/11 - - A + 2/8%F304H 0.04 - 0.10 0.75 2.0 18/20 8/11 - - A + 2/8%F304N 0.08 0.75 2.0 18/20 8/11 - 0.1/0.16N A + 2/8%F316 0.08 0.75 2.0 16/18 11/142/3 - A + 3/10%F347 0.08 0.75 2.0 17/20 9/13 - Nb : 10xC A + 4/12%F321 0.08 0.75 2.0 17/19 9/12 - Ti: 5xC A + 4/12%F310 0.15 0.75 2.0 24/26 19/22 - - 100% A309 0.08 1.0 2.0 22/24 12/15 - - A + 8/15%F308L (generally filler metal only) 0.03 1.0 2.0 19/21 10/12     A + 4/12%FThe 3XX may followed by a letter that gives more information about the specific alloy as shown in the Table. 'L' is for a low carbon austenitic stainless steel for use in an aggressive corrosive environment ; 'H' for a high carbon steel with improved high temperature strength for use in creep applications; 'N' for a nitrogen bearing steel where a higher tensile strength than a conventional steel is required. These suffixes are used with most of the alloy designations egtype 316L, type 316LN, type 347H, where the composition has been modified from that of the base alloy.

Austenitic stainless steels are metalurgically simple alloys. They are either 100% austenite or austenite with a small amount of ferrite (see Table 1). This is not the ferrite to be found in carbon steel but a high temperature form known as delta (δ) -ferrite. Unlike carbon and low alloy steels the austenitic stainless steels undergo no phase changes as they cool from high temperatures. They cannot therefore be quench hardened to form martensite and their mechanical properties to a great extent are unaffected by welding. Cold (hydrogen induced) cracking (Job Knowledge No. 45) is therefore not a problem and preheat is not necessary irrespective of component thickness.

Alloying elements in an austenitic stainless steel can be divided into two groups; those that promote the formation of austenite and those that favour the formation of ferrite. The main austenite formers are nickel, carbon, manganese and nitrogen; the important ferrite formers are chromium, silicon, molybdenum and niobium. By varying the amounts of these elements, the steel can be made to be fully austenitic or can be designed to contain a small amount of ferrite; the importance of this will be discussed later.

In 1949 Anton Schaeffler published a constitutional or phase diagram that illustrates the effects of composition on the microstructure. In the diagram Schaeffler assigned a factor to the various elements, the factor reflecting the strength of the effect on the formation of ferrite or austenite; these factors can be seen in the diagram. The elements are then combined into two groups to give chromium and nickel 'equivalents'. These form the x and y axes of the diagram and, knowing the composition of an austenitic stainless steel, enables the proportions of the phases to be determined.

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Fig 1. Shaeffler diagram (A-austenite; M - martensite; F - ferrite)

Typical positions of some of the commoner alloys are given in Fig.1. Also superimposed on this diagram are coloured areas identifying some of the fabrication problems that may be encountered with austenitic stainless steels.

Although all the austenitic stainless steels are sensitive to hot cracking (Job Knowledge No.44), the fully austenitic steels falling within the vertically blue area in Fig.1 such as type 310 are particularly sensitive. 

The main culprits are sulphur and phosphorus. To this end, these tramp elements have been progressively reduced such that steels with less than 0.010% sulphur and phosphorus less than 0.020% are now readily available. Ideally a type 310 or type 317 alloy should have sulphur and phosphorus levels below some 0.003%. Cleanliness is also most important and thorough degreasing must be carried out immediately prior to welding.

The steels such as type 304, type 316, type 347 that fall within, or close to, the small uncoloured triangular region in the centre of the diagram contain a small amount of delta-ferrite and, whilst not being immune to hot cracking, have improved resistance to the formation of sulphur-containing liquid films. The reasons for this are that a) ferrite can dissolve more sulphur and phosphorus than austenite so they are retained in solution rather than being available to form liquid films along the grain boundaries and b) the presence of quite a small amount of ferrite increases the grain boundary area such that any liquid films must spread over a greater area and can no longer form a continuous liquid film. The 100% austenitic steels do not have this advantage.

One problem that has arisen with very low sulphur steels is a phenomenon known as 'cast to cast variation' or 'variable penetration'. The weld pool in a low sulphur steel (<0.005%) tends to be wide with shallow penetration; a steel with sulphur over some 0.010% has a narrower, more deeply penetrating weld bead.

This is generally only a problem with the use of the fully automated TIG welding process, a manual welder being capable of coping with the variations in penetration due to the differences in sulphur content in different casts of steel. However, automated TIG welding procedures developed on a 'high' sulphur steel, when used to weld a low sulphur steel may result in lack of penetration type defects; the reverse situation may result in excessive penetration.

Changes to the procedure that have mitigated, but never eliminated, this problem have included slow travel speed, pulsed current, use of Ar/H2 shield gas mixtures. Other methods include specifying a minimum sulphur of, say, 0.010% or segregating the steels into batches with known penetration characteristics and developing welding procedures to suit. The A-TIG activated flux process has also been found to be of benefit.

The problems of welding the fully ferritic steels that fall into the pink area, where grain growth and embrittlement is a problem, have already been dealt with in Job Knowledge No.101.

The austenitic stainless steels falling into the yellow area will also embrittle but this is as a result of the formation of hard brittle phases called 'sigma' (σ) and 'chi' (χ). This embrittlement takes place in the temperature range of approximately 500 to 900°C. It is a sluggish process and is not a problem during welding of the austenitic stainless steels, but can occur in elevated temperature service or if the welded component is stress relieved.

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Formation of these phases is promoted by high chromium and molybdenum (ferrite forming elements) so that steels such as type 310 and type 316 are particularly sensitive and may show a substantial loss of ductility after stress relief. Delta ferrite also transforms more rapidly that austenite so those alloys containing large amounts of this phase will degrade faster than an austenitic steel with only a small percentage of ferrite; hence the problems with duplex and super duplex stainless steels.

When it is necessary to stress relieve a fabrication then the loss of ductility must be accounted for. In those steels containing delta-ferrite this phase should be held to a minimum, consistent with minimising the risk of hot cracking, by control of the ferrite forming elements by requiring typically 2% to 5% delta ferrite.

The next Job Knowledge article will look at weld filler metal selection, some of the service problems of the austenitic stainless steels and how these may be mitigated.

Welding of austenitic stainless steel. Part 2

Job Knowledge

The previous Job Knowledge article, No. 103, dealt with the metallurgy of austenitic stainless steels and some of the welding problems that may be encountered.

Austenitic stainless steels can be welded with all the commercially available welding processes. There are matching filler metals available for most of the austenitic range of alloys, the exceptions being that there is no type 304 filler metal available (this alloy is generally welded with type 308 filler metal) and no type 321 filler due to the problems of transferring titanium across the arc. Type 321 steels are usually welded with a type 347 filler.

Also mentioned In Job Knowledge 103 was that the austenitic stainless steels are metallurgically simple alloys and room temperature mechanical properties are not significantly affected by variations in the welding procedure. However, increasing the oxygen and ferrite levels will reduce the toughness at cryogenic (~-196°C) temperatures.

Basic coated manual metal arc electrodes with a controlled short arc length and basic agglomerated submerged arc fluxes are required for best toughness if arc welding processes are used. The steel and filler metal should be selected with as low a ferrite content as possible, say 1 to 3% for best Charpy-V test results.

Conversely, for best creep resistance an 'H' grade steel should be selected and rutile or acid/rutile electrodes and acid submerged arc fluxes should be used. These improve the creep strength by increasing the titanium and niobium content of the weld metal, forming a greater concentration of grain strengthening carbides.

TIG (GTAW) welding of the root pass must always be carried out with an inert gas back purge to prevent loss of chromium (and hence of corrosion resistance), argon being the gas generally used for this purpose. Nitrogen may be used but there is a risk of the weld deposit absorbing nitrogen, thereby becoming fully austenitic and hot crack sensitive.

Two characteristics of austenitic stainless steels that differentiate them from ferritic steels are the coefficients of thermal conductivity and expansion. Austenitic stainless steels have a low coefficient of thermal conductivity, approximately 1/3rd that of ferritic steel at room temperature and a coefficient of thermal expansion some 30% more than that of a ferritic steel.

Higher expansions in a narrower HAZ result in higher residual stresses and more distortion. This is a particular problem with thin sheet fabrications where the achievement of the desired dimensional tolerances can be extremely difficult and costly to achieve. The use of accelerated cooling techniques such as copper

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chills or a freezing gas (the liquid CO2low stress-no distortion technique typifies this approach) have been used to reduce distortion to acceptable levels.

One of the main reasons for using an austenitic stainless steel is its corrosion resistance. Whilst this is primarily a function of the chromium content of the steel, carbon also has a major but adverse effect resulting in a form of corrosion known as intergranular or intercrystalline corrosion (ICC) or weld decay, a localised effect confined to the HAZ.

Carbides present in the HAZ of an austenitic stainless steel dissolve on heating and reform on cooling during the welding heat cycle. Unfortunately, these new precipitates form preferentially as chromium carbides on the grain boundaries, depleting chromium from the region immediately adjacent to the boundary, resulting in a local loss of chromium and a reduction in corrosion resistance. If sufficient chromium carbides are formed this results in a network of steel along the grain boundaries sensitive to corrosion; the steel has been sensitised. This sensitisation occurs in the HAZ region that has seen temperatures between 600 and 900°C and times that may be as short as 50 seconds.

There are several methods that may be used to overcome this difficulty. A solution heat treatment (1050°C followed by a water quench) will re-dissolve the carbides and these will be retained in solution on rapid cooling. Whilst this will eliminate the chromium depleted regions it is rarely practical to solution-treat complex welded structures.

The most obvious alternative technique is to reduce the carbon content. This has two beneficial effects:

The lower the carbon content, the longer the time required to form the carbides. At 0.08% carbon this time is around 50 seconds; at 0.03% carbon the time required is about eight hours, most unlikely to be achieved during welding!

The lower the carbon content then the fewer carbides there are to form a continuous chromium depleted network. Hence the 'L' grades, type 304L, or 316L, are preferred where best corrosion resistance is required.

One other method is the addition of alloying elements that will form carbides in preference to chromium; thus the stabilised type 321 and 347 grades containing titanium and niobium respectively were developed.

Titanium and niobium are very strong carbide formers that precipitate carbides at higher temperatures than those at which chromium carbides will form so there is no carbon available to react with the chromium. However, even these stabilised grades may corrode in a very narrow band close to the fusion line (the so-called knife-line attack) in the presence of hot acids. This is due to the higher and more restricted temperature range at which the niobium or titanium carbides dissolve. The solution, as above, is to limit the carbon to 0.03% maximum.

Welding consumables must also be selected with low carbon content if best corrosion resistance is required. Most arc welding consumables contain less than 0.03% carbon but there are filler metals available with carbon contents of up to 0.10%; these should only be used to weld the 'H' grades of steel where good creep resistance is required.

Although MAG (GMAW) welding is often used it should be remembered that carbon pick-up is possible when argon/CO2mixtures are used, particularly if the welding is carried out in the dip transfer mode. Argon/2% oxygen mixtures are therefore generally preferred where best corrosion resistance is required but argon/10% CO2/2% oxygen is a good compromise that can be used for a broad range of applications.

The other major service problem encountered with the austenitic stainless steels is that of stress corrosion cracking. This may be caused by strong alkali solutions but it is the halides (chlorides, fluorides and bromides) that are primarily responsible. Cracking takes place in areas of high stress, as the name suggests,

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and is not therefore confined solely to welds, but it is at and adjacent to welds that stresses approaching the yield point of the metal are found and these present a particular problem.

The cracking is transgranular and propagation rates can be extremely rapid given the ideal conditions. In hot concentrated chloride solutions, for example, penetration can occur in thin, sheet components within a few minutes. However, the lower the temperature and/or the acid concentration then the rate of crack propagation is correspondingly slower. Austenitic stainless steels are therefore not generally used where halides are present. Even here, stress corrosion cracking (SCC)may occur due to contamination, either of the product in the pipe or vessel or externally from sea water, particularly where the liquid is able to concentrate in crevices.

To eliminate any chance of SCC, the only solution is to stress relieve the weld at a temperature of around 700 to 900°C. It should be remembered that:

this may sensitise the steel so only low carbon grades should be used and the steel may embrittle due to sigma phase formation (see Job Knowledge 103) at the lower heat

treatment temperatures.

Stress corrosion cracking (SCC) in Type 316L stainless steel

Local stress relief should be approached with caution as the temperature gradients may result in stresses developing outside the heated band; wider heated bands and more stringent control of temperature gradients than required by specifications or codes may therefore be necessary. Solution treatment (1050°C soak followed by very rapid cooling, ideally a water quench) will eliminate all residual stresses whilst avoiding both sensitisation and embrittlement but is rarely practical on a welded assembly.

The alternative is to select a steel that is more resistant; the molybdenum bearing grade type 316 is better than 304 or 321. The ferritic stainless steels (Job Knowledge 101) are not susceptible to chloride SCC.

Welding of ferritic/martensitic stainless steels

Job Knowledge

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Stainless steels are 'stainless' i.e. are corrosion resistant, due to the presence of chromium in amounts greater than 12%, where it forms a passive film on the surface of the steel. Note that these stainless steels are not the 'stainless steels' that generally first spring to mind; the 18% Cr/8% Ni austenitic stainless steels of the Type 304 or Type 316 grades; but two separate groups of alloys with different mechanical and corrosion resistant properties.

The ferritic stainless steels contain up to some 27% chromium and are used in applications where good corrosion/oxidation resistance is required but in service loads are not excessive, e.g. flue gas ducting, vehicle exhausts, road and rail vehicles.

The martensitic grades contain up to 18% chromium and have better weldability and higher strengths than the ferritic grades. They are often found in creep service and in the oil and gas industries where they have good erosion and corrosion resistance.

Now for a little metallurgy! Chromium is an alloying element that promotes the formation of ferrite in steel; in the case of the ferritic stainless steels, this ferrite is the high temperature form known as delta-ferrite. Unlike the low alloy steels, therefore, this type of steel undergoes no phase changes as it cools from melting point down to room temperature; they cannot therefore be hardened by heat treatment and this has implications with respect to the properties of welded joints.

Carbon and nitrogen, however, are two elements that promote the formation of austenite so, as the percentage of carbon and/or nitrogen increases, the ferritic steel can be designed to transform, wholly or partially, to austenite before transforming back to ferrite. This series of phase changes are similar to those in a low alloy steel, enabling the steel to be hardened by producing martensite - the martensitic stainless steels. Compositions and typical properties of some of the alloys are given in Table 1.

Table 1 Typical properties of ferritic and martensitic steels

AISI Number Steel TypeChemical Composition (max %) Mechanical Properties

(annealed cond; typical)C Mn Cr Ni Mo UTS (MPa) Y.S. (MPa) El.%

409 ferritic 0.08 1.0010.5/11.75 - - 480 240 25430 ferritic 0.12 1.0016.0/18.0     520 345 25434 ferritic 0.12 1.0016.0/18.0   0.75/1.25 530 370 22446 ferritic 0.20 1.5 23.0/27.0     550 350 20410 martensitic 0.15 1.0011.5/13.00 - - 480 310 25420(API 5CT L-80)

martensitic 0.15 min 1.0012.0/14.0 - - 650 345 25

422(12CrMoV) martensitic 0.25 1.3 10.0/12.0 0.8 1.2 (V 0.4) 720 550 22

431 martensitic 0.20 1.0015.0/17.0 1.25/2.5   860 670 20There are a number of welding problems with the ferritic steels. Although they are not regarded as hardenable, small amounts of martensite can form, resulting in a loss of ductility. In addition, if the steel is heated to a sufficiently high temperature, very rapid grain growth can occur, also resulting in a loss of ductility and toughness.

Although the ferritic steels contain only small amounts of carbon, on rapid cooling carbide precipitation at the grain boundaries can 'sensitise' the steel making it susceptible to inter-crystalline corrosion. When this is

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associated with a weld it is often known as weld decay. Developments in recent years of extra low carbon, titanium or niobium containing grades have, however, improved this situation.

The ferritic stainless steels are generally welded in thin sections. Most are less than 6mm in thickness where any loss of toughness is less significant. Most of the common arc welding processes are used although it is regarded as good practice to limit heat input with these steels to minimise grain growth (1kj/mm heat input and a maximum interpass temperature of 100-120°C is recommended) implying that the high deposition rate processes are inadvisable. Preheat is not required although it may be helpful when welding sections over, say, 10mm thick, where grain growth and welding restraint may result in cracking of the joint.

Welding consumables for the ferritic steels are generally of the austenitic type; type 309L (low carbon grade) is the most commonly used. This is to ensure that any dilution that occurs does not result in a low ductility austenitic/ferritic/martensitic weld metal micro-structure. However, provided care is taken to control dilution, types 308 and 316 may be used. Nickel based consumables may also be used and will result in better service performance where the component is thermally cycled. A matching filler metal is available for welding of Grade 409 steel, often used in vehicle exhaust systems.

Post weld heat treatment (PWHT) at around 620°C is rarely carried out although a reduction in residual stress will give an improved fatigue performance: nickel based fillers are a better choice in this context than the Cr/Ni austenitic consumables.

The martensitic grades are used in more challenging environments and, as the name suggests, present rather more problems than the ferritic steels. Both the higher carbon (>0.1%) and low carbon (<0.1%) versions, with a few exceptions, require preheat and PWHT to avoid weldment cracking problems and to provide a sufficiently tough and ductile joint.

Matching welding consumables are available for most grades so that corrosion resistance and mechanical properties can be matched to those of the parent metal. To reduce the risk of hydrogen induced cracking, low hydrogen welding processes are essential and preheat temperatures of 200 to 300°C are recommended. A weld that has been completely transformed to untempered martensite by allowing the joint to cool to room temperature can be extremely brittle and great care is needed in handling to prevent brittle failure. In addition, such joints are sensitive to stress corrosion cracking even in a normal fabrication shop environment. It is highly advisable therefore to PWHT as soon as possible on completion of welding.

A conventional heat treatment cycle would be to cool the joint to below 100°C to ensure full transformation of the weld and HAZ to martensite, closely controlled heating to minimise stresses from temperature variations, PWHT at around 700°C for one to four hours and controlled cool to ambient.

A hydrogen release treatment from the preheat temperature, say 350°C for four hours, is unlikely to reduce the risk of cold cracking. If the steel is not allowed to cool to a sufficiently low temperature so that full transformation to martensite takes place then there will be austenite present during the hydrogen release treatment.

This austenite will retain hydrogen and may generate cracks when it transforms to martensite as the joint is cooled to ambient. If cold cracking is a real issue, even with good hydrogen control, then it may be necessary to PWHT directly from the preheat temperature, cool to ambient and repeat the PWHT to temper any martensite that was formed following the first cycle of PWHT.

Welding consumables matching the base metal composition are available for most of the martensitic stainless steels, often with small additions of nickel to ensure that no ferrite is formed in the weld. Nickel lowers the temperature at which martensite transforms to austenite so it is important with such filler metals

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that the PWHT temperature is not allowed to exceed about 750°C otherwise untempered martensite will form in the weld as the item cools to ambient.

Conventionally, when welding dissimilar metal joints the filler metal is selected to match the composition of the lower alloyed steel. Experience has shown that this can cause cold cracking problems so filler metals matching the martensitic steel should be used. An alternative is to weld with austenitic stainless steel fillers, type 309 for example, but the weld may then not match the tensile strength of the ferritic steel and this must be recognised in the design of the weld. Nickel based alloys may also be used; alloy 625 for instance, has a 0.2% proof strength of around 450MPa; and will give a better match on coefficient of thermal expansion.

The metallurgy of these types of steels is complex and they are frequently used in challenging and safety related environments. An article such as this can only give a partial picture so if there are any doubts surrounding their fabrication it is recommended that advice is sought from suitable specialists.

Duplex stainless steel. Part 1

Job Knowledge

The name 'duplex' for this family of stainless steels derives from the microstructure of the alloys which comprises approximately 50/50 mixture of austenite and delta-ferrite. They are designed to provide better corrosion resistance, particularly chloride stress corrosion and chloride pitting corrosion, and higher strength than standard austenitic stainless steels such as Type 304 or 316. The main differences in composition, when compared with an austenitic stainless steel is that the duplex steels have a higher chromium content, 20 - 28%; higher molybdenum, up to 5%; lower nickel, up to 9% and 0.05 - 0.5% nitrogen. Both the low nickel content and the high strength (enabling thinner sections to be used) give significant cost benefits. They are therefore used extensively in the offshore oil and gas industry for pipework systems, manifolds, risers, etcand in the petrochemical industry in the form of pipelines and pressure vessels.

In addition to the improved corrosion resistance compared with the 300 series stainless steels duplex steels also have higher strength. For example, a Type 304 stainless steel has a 0.2% proof strength in the region of 280N/mm2, a 22%Cr duplex stainless steel a minimum 0.2% proof strength of some 450N/mm2 and a superduplex grade a minimum of 550N/mm2.

Although duplex stainless steels are highly corrosion and oxidation resistant they cannot be used at elevated temperatures. This is due to the formation of brittle phases in the ferrite at relatively low temperatures, see below, these phases having a catastrophic effect on the toughness of the steels. The ASME pressure vessel codes therefore restrict the service temperature of all grades to below 315°C, other codes specify even lower service temperatures, perhaps as low as 250°C for superduplex steels.

Duplex alloys can be divided into three main groups; lean duplex, 22%Cr duplex and 25%Cr superduplex, and even higher alloyed, hyperduplex grades have been developed, this division being based primarily on the alloy's alloying level, eg in terms of 'PREN' (pitting resistance equivalence number), a measure of the

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alloy's resistance to pitting corrosion. PREN is calculated from a simple formula: PREN = %Cr + 3.3%Mo +16%N and an allowance for W is sometimes made, having a factor of 1.65. A duplex steel has a PREN less than 40; a superduplex a PREN between 40 and 45 and hyperduplex a PREN above 45, whilst the lean grades typically have lower nickel and hence lower price.

The commonest shorthand method of identifying the individual alloys is by the use of the trade name, particularly for the superduplex grades, eg UR52N+, Zeron 100, 2507 or DP3W, whilst the most common 22%Cr grade, UNS S31803 has widely become known as 2205 regardless of its supplier, although this is a trade name.

The UNS numbering system offers an independent alternative. Typical compositions and minimum proof strengths of the more common duplex alloys are given in the Table. Note that the commonly used 2205 applies to two UNS numbers, S31803 and S32205, with S32205 being a more recent and controlled composition.

Typical compositions and proof strengths of common duplex stainless steels

CommonName UNS No BS EN

NoSteelType

Typical Chemical Composition % 0.2%proof

strengthN/mm2 (min)

%C Cr Ni Mo N Cu

2304 S32304 1.4362 duplex 0.015 23.0 4.0 0.055 0.13   4002205 S31803 1.4462 duplex 0.015 22.0 5.5 3.0 0.14 - 4502205 S32205 1.4462 duplex 0.015 22.5 5.5 3.3 0.17   450255(UR52N) S32520 1.4507 super duplex 0.015 25.0 7.0 3-5 0.28 0.13 5502507 S32750 1.4410 super duplex 0.015 25.0 7.0 4.5 0.28 0.3 550Zeron 100 S32760 1.4501 super duplex 0.015 25.0 7.0 3.5 0.25 0.8 550Sandvik SAF3207 S33207 - hyper duplex 0.03 31 7.5 4.0 0.50 0.75 700The metallurgy of the duplex stainless steel family is complex and requires very close control of composition and heat treatment regimes if mechanical properties and/or corrosion resistance are not to be adversely affected. To produce the optimum mechanical properties and corrosion resistance the microstructure or phase balance of both the parent and weld metal should be around 50% ferrite and 50% austenite. This precise value is impossible to achieve repeatably but a range of phase balance is acceptable. The phase balance of parent metals generally ranges from 35 - 60% ferrite.

Whilst composition and, perhaps more importantly, heat treatment parameters are relatively easy to control this is not the case during welding. The amount of ferrite is dependant not only on composition but also on the cooling rate; fast cooling rates retain more of the ferrite that forms at elevated temperature. Therefore to minimise the risk of producing very high ferrite levels in the weld metal it is necessary to ensure that there is a minimum heat input and therefore a maximum cooling rate. A rule of thumb is that heat input for duplex and superduplex steels should be not less than 0.5kJ/mm although thick sections will need this lower limit to be increased.

Welding consumables are also generally formulated to contain more nickel than the parent metal, nickel being one of the elements that promotes the formation of austenite. A duplex filler metal may contain up to 7% nickel, a superduplex up to 10% nickel.

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Reference to the phase diagrams and CCT curves shows that the duplex stainless steels fall within the area where the production of brittle intermetallic phases is a major risk during welding and heat treatment, markedly reducing both toughness and corrosion resistance.

The main culprits are sigma phase, chi phase and 475°C embrittlement. Sigma and chi phases form at temperatures between 550 and 1000°C with the fastest rate of formation around 850°C. The time to form these phases can be as short as 30 or 40 seconds in a superduplex alloy. 4750C embrittlement, as the name suggests, occurs at lower temperatures of some 350 - 550°C with times for the start of formation of perhaps 7 - 10 minutes.

Short times such as these are within the ranges that may be encountered during interpass cooling so, once again, heat input and cooling rates become very important welding parameters except that this time it is the maximum heat input that needs to be controlled. A maximum heat input of 2.5kJ/mm should be acceptable for the duplex steels and 2.0kJ/mm maximum for superduplex. Many codes and contract specifications, however, further restrict heat inputs to less than 1.75 - 2kJ/mm for duplex steels and 1.5 - 1.75kJ/mm for superduplex.

Two other factors that also affect cooling rates are preheating and interpass temperatures. Preheat is not generally regarded as necessary for duplex stainless steels unless the ambient conditions mean that the steel is below 5°C or there is condensation on the surface. In these situations a preheat of around 50 - 75°C should be adequate. Very thick section joints, particularly those welded with the submerged arc process, can also benefit from a low preheat of around 100°C.

Interpass temperature can have a significant effect on the microstructure of the weld and its heat affected zones. For a duplex steel 250°C is regarded as an acceptable maximum and for a superduplex 150°C maximum. Note, however, that many codes do not separate the grades into duplex and superduplex and 150°C is often required as the norm. Such low interpass temperatures can have a serious effect on joint completion times and forced cooling by blowing dry air through the bore of a pipe once the bore purge has been removed has been used. This is generally only beneficial when thick wall vessels or pipes are being welded using a rotated pipe mechanised TIG process or submerged arc. If this technique is used then it is advisable to force cool the procedure qualification test piece to ensure that cooling rates (and the resultant microstructures) are within the permissible range.

Care therefore needs to be taken to read through code and contract specification requirements and to ensure that the requirements with respect to heat input, interpass temperature etc. are incorporated in welding procedure documentation prior to welding procedure qualification. The next Job Knowledge will provide some guidelines for the welding of the duplex stainless steels.

Duplex stainless steel - Part 2

Job Knowledge

Part 1

The previous article highlighted some of the problems encountered when welding duplex and superduplex stainless steels, in particular the need to control closely the heat input if an undesirable phase balance or the formation of brittle intermetallic phases are to be avoided.

This requirement has implications with respect to quality control. Variations in weld preparations which would be compensated for by the welder changing his welding technique, wide root gaps for example, may result in a significant change in heat input. Weld preparations therefore need to be more closely controlled than for a conventional stainless steel.

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It is recommended that weld preparations are machined for greatest accuracy but, if hand-ground, close attention must be paid to the weld preparation dimensions. Welding supervisors and inspectors also need to understand the importance of heat input control, ensuring that welding is not allowed to take place outside the limits of the qualified procedures with regular checking of welding parameters and interpass temperature.

Hot cracking is rarely a problem due to the high ferrite content but has been observed, particularly in submerged arc welds. Cleanliness of the joint is therefore still important. Machining or grinding burrs and any paint should be removed and the joint thoroughly degreased and dried prior to welding. Failure to do so can affect corrosion resistance and joint integrity.

Hydrogen cold cracking, whilst unusual, is not unknown and can occur in the ferrite of weld metal and HAZs at quite low hydrogen concentrations. It is recommended that the hydrogen control measures used for low alloy steel consumables should apply for duplex consumables. Submerged arc fluxes and basic coated electrodes should be baked and used in accordance with the manufacturer's recommendations; shield gases must be dry and free of contaminants.

Most commercially available welding consumables will provide weld metal with yield and ultimate tensile strengths exceeding those of the parent metal but there is often difficulty in matching the notch toughness (Charpy V) values of the wrought and solution treated base metal.

TIG welding gives very clean weld metal with good strength and toughness. Mechanisation has substantially increased the efficiency of the process such that it has been used in applications such as cross-country pipelining.

Gas shielding is generally pure argon although argon/helium mixtures have given some improvements by permitting faster travel speeds. Nitrogen, a strong austenite former, is an important alloying element, particularly in the super/hyper duplex steels and around 1 to 2% nitrogen is sometimes added to the shield gas to compensate for any loss of nitrogen from the weld pool. Nitrogen additions will, however, increase the speed of erosion of the tungsten electrode. Purging the back face of a joint is essential when depositing a TIG root pass. For at least the first couple of fill passes pure argon is generally used although small amounts of nitrogen may be added and pure nitrogen has occasionally been used.

TIG welding may be performed without any filler metal being added but is not recommended on duplex steels as the corrosion resistance will be seriously impaired. Filler metals are be selected to match the composition of the parent metal but with an additional 2 to 4% nickel to ensure that sufficient austenite is formed. Any stray arc strikes will be autogenous and must be removed by grinding.

MMA welding is carried out with matching composition electrodes overalloyed with nickel and either rutile or basic flux coatings. Basic electrodes give better notch toughness values. Electrodes of up to 5mm diameter are available with the smaller diameters providing the best control when welding positionally.

MAG welding is generally carried out using wires of 0.8 to 1.2mm diameter, rarely exceeding 1.6mm and of a similar composition to the TIG wires. Shielding gases are based on high purity argon with additions of carbon dioxide or oxygen, helium and perhaps nitrogen. Because of the presence of carbon dioxide or oxygen the weld metal notch toughness (Charpy V values) are less than can be achieved using TIG. Microprocessor-controlled pulsed welding gives the best combination of mechanical properties. Mechanisation of the process is easy and can give significant productivity improvements although joint completion times may not be as short as anticipated due to the need to control interpass temperatures to below the recommended maximum.

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Flux-cored arc welding (FCAW) is used extensively with major productivity gains being possible in both manual and mechanised applications. The flux core is generally rutile; the shielding gas CO2, argon/20%CO2 or argon/2%O2. The presence of carbon dioxide or oxygen leads to oxygen, and, in the case of CO2, carbon pickup in the weld metal, thus notch toughness is reduced. Metal cored wires are also available that require no slag removal; better suited to mechanised applications than flux-cored wires. Because of differences in flux formulation and wire composition between manufacturers it is recommended that procedure qualification is carried out using the specific make of wire used in production even though the wires may fall within the same specification classification.

Submerged arc welding (SAW) is generally confined to welding thick wall pipes and pressure vessels. Solid wires, similar to those available for TIG welding, are available. Fluxes are generally acid-rutile or basic, the latter giving the best toughness values in the weld metal. As with any continuous mechanised welding process the interpass temperature can rapidly increase and care needs to be taken to control both interpass temperature and process heat input. Because of the need to control heat input the wire diameter is normally limited to 3.2mm permitting a maximum welding current of 500A at 32V although larger diameter wires are available. However, any productivity gains from the use of a large diameter wire and high welding current may not be realised due to the need for interpass cooling.

There is often the need to weld duplex/superduplex steel to lower alloyed ferritic steel, a 300 series stainless steel or a dissimilar grade of duplex steel. The 300 series stainless steels are generally welded to duplex steels with a 309MoL (23Cr/13Ni/2.5Mo) filler metal. Low carbon and low alloy steels may be welded to duplex steels using either a 309L (23Cr/13Ni) or a 309MoL filler metal.

These two filler metals, however, have yield and ultimate tensile strengths substantially less than most low carbon/low alloy steels and all duplex steels. This means the designer has to take this reduction of strength into account by increasing the component thickness or the welding engineer has to select a filler metal that both matches the strength of the weaker steel and is compatible with the two parent metals. These considerations narrow the choice to one of the nickel-based alloys such as alloy 82 or, for higher strength, a niobium-free high alloyed nickel filler, such as C22. or 59. Alloy 625 has been used but problems with reduced toughness due to the formation of niobium nitride precipitates along the fusion boundary have resulted in the alloy falling out of favour.

Duplex steel welds are seldom post-weld heat treated. Due to sigma phase formation they cannot be given a heat treatment at the low temperatures of 600-700°C, the normal range for stress relief unless a qualification programme has been undertaken to demonstrate that the loss of toughness is acceptable. If PWHT is required then ideally the whole component must be given a solution anneal at 1000-1100°C followed by a water quench; an impractical operation with most welded structures.

Lastly, any process that heats the steels above 300°C will affect the mechanical properties. Heat straightening to control distortion should therefore not be carried out. The HAZs produced by hot cutting processes like plasma or laser may contain undesirable microstructures. Cut edges that will enter service 'as-

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cut' must be ground or machined back for a minimum of 2mm to remove the HAZ and ensure there is no loss of toughness or corrosion resistance.

If the cut edges are welded after cutting then the HAZs are generally sufficiently narrow that the effects of the cutting operation are lost although it is recommended that, as above, the edges are ground or machined back 2mm.

Precipitation hardening stainless steels

Job Knowledge

The precipitation hardening (PH) stainless steels are a family of corrosion resistant alloys some of which can be heat treated to provide tensile strengths of 850MPa to 1700MPa and yield strengths of 520MPA to over 1500MPa - some three or four times that of an austenitic stainless steel such as type 304 or type 316. They are used in the oil and gas, nuclear and aerospace industries where a combination of high strength, corrosion resistance and a generally low but acceptable degree of toughness is required. Precipitation hardening is achieved by the addition of copper, molybdenum, aluminium and titanium either singly or in combination.

The family of precipitation hardening stainless steels can be divided into three main types - low carbon martensitic, semi-austenitic and austenitic - typical compositions of some of the steels are given in Table 1.

Table 1 Typical Compositions of some commoner precipitation hardening stainless steels

Specification Common Name TypeTypical Chemical Analysis %

C Mn Cr Ni Mo Cu Al Ti OthersA693 Tp630 17/4PH martensitic 0.050.7516.5 4.25 - 4.25 - - Nb 0.3  FV 520 austenitic-martensitic 0.050.6 14.5 4.751.4 1.7 - - Nb 0.3A693 Tp631 17/7PH austenitic-martensitic 0.060.7 17.257.25 - - 1.25 - -  PH 15/7 Mo austenitic-martensitic 0.060.7 15.5 7.252.6 - 1.3 - -

A 286   austenitic 0.041.4515.2526.01.25 - 0.152.15V 0.25B 0.007

  JBK 75 austenitic 0.010.0414.7530.51.25 - 0.302.15V 0.25B 0.0017

  17/10P austenitic 0.070.7517.2 10.8         P 0.28The martensitic PH steels, of which 17/4PH is the most common, transform to martensite at low temperatures, typically around 250°C, and are further strengthened by ageing at between 480 and 620°C.

The austenitic-martensitic PH steels are essentially fully austenitic after solution treatment and require a second heat cycle to 750°C/2 hours before cooling to room temperature to form martensite. Some of these alloys need to be refrigerated (-50/-60°C for eight hours) following this heat treatment to ensure full transformation to a stable austenitic/martensitic structure although the two most commonly used alloys, FV520 and 17/7PH, do not require refrigeration to develop optimum properties.

Ageing of these alloys occurs at temperatures between 500 to 600°C. The austenitic grades are stable down to room temperature, improvements in strength being from the precipitates formed by ageing at 650 to 750°C. These fully austenitic grades can exhibit good toughness and some may be used at cryogenic temperatures.

For best weldability it is recommended that all three types of alloys are supplied in the annealed, solution treated or overaged condition. Alloys in the form of sheet or strip may be in a cold worked condition and

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weldability is seriously compromised. As with many precipitation hardening alloys, achieving mechanical properties in the weld and HAZs to match those of the parent material is a problem. Even with matching welding consumables, a full solution treatment and age hardening the maximum strength of a joint in the semi-austenitic and austenitic alloys is likely to be only some 90% of that of the base metal.

Martensitic PH steels in the solution-treated condition can be welded with most of the conventional arc welding processes although the best toughness will be achieved with the TIG (GTAW) process as this provides the cleanest weld metal. Even better toughness can be achieved using power beam processes (electron beam or laser welding). Matching filler metals are available for most of the steels in this group enabling matching mechanical properties to be achieved by carrying out a post weld ageing heat treatment.

If a joint is very highly restrained then 17/4PH may fail along the fusion line by a form of reheat cracking during the ageing heat treatment. In these circumstances the component should be welded in the overaged condition and then given a solution heat treatment followed by the PWHT described below. Austenitic filler metals such as 308L or, for higher weld metal strength, a duplex filler metal such as 2205, can be used where lower strength joints can be tolerated or cracking due to high restraint is a problem. PWHT is not possible if a duplex filler metal is used or recommended for austenitic weld metal due to embrittlement.

The martensite in these steels is relatively soft due to the low carbon content so preheat is not generally necessary although for thick, (above 25mm) highly restrained joints, a preheat of around 100°C has been found to be useful in reducing the risk of cracking. Because of the low temperature at which these steels transform to martensite a maximum interpass temperature of 200°C is recommended.

Maintaining a very high interpass temperature results in the entire weld transforming to martensite on cooling to room temperature and the volume change that occurs when this happens can then lead to a form of quench cracking. The stress raising effect of the notch in the root of fillet welds and partial penetration butt welds has been found to cause cracking. Provided the reduction in strength can be tolerated, a Tp308L root pass can be used to solve this problem. It has also been found that 17/4PH castings may form HAZ hot cracks during welding; for cast items the copper content is therefore limited to 3% maximum.

PWHT generally comprises a 750°C soak and cool to room temperature to ensure that the steel is 100% martensitic followed by ageing at 550°C. This should give UTS of 900 to 1000MPa, yield strength 800 to 900MPa and ductility of some 15% depending upon the composition of the alloy and the temperature of the ageing heat treatment.

The semi-austenitic alloys are generally supplied in the solution treated condition. This means that the steel is fully austenitic and preheat is not generally required although for welding of thick and highly restrained joints a preheat of around 100°C has been found to be helpful. All the common arc welding processes may be used although, as above, TIG (GTAW) will give the best properties.

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For alloys containing aluminium, eg 17/7PH, MMA and submerged arc welding should be avoided as a good proportion of the aluminium is lost during welding; inert gas shielded processes are therefore preferred. The weld pool is less fluid than the non-aluminium alloys. Matching composition filler metals for FV520 are readily available but 17/7PH consumables are difficult and expensive to obtain so parent metal sheared from strip is often used for TIG welding. Alternatively a 17/4PH or FV520 filler may be used; a preheat of 100°C is advisable if the 17/4PH filler is used. PWHTs are similar to those used for the martensitic steels but, without a full solution heat treatment and matching filler metal, strengths matching those of the parent metal are unlikely to be achieved.

It is recommended that the fully austenitic PH steels are welded in the solution treated condition; a water or oil quench from around 980°C. The ageing process is very sluggish, requiring some 15 hours at 720°C to develop full strength and this means that the HAZ is virtually unchanged from the parent metal. Optimum strength can therefore be developed during the post-weld ageing treatment. These steels, like the austenitic stainless steels, are insensitive to cold cracking and do not require to be pre-heated. They are, however, very sensitive to hot cracking due to them being fully austenitic. This makes the welding of thick sections problematic and requires the welding conditions to be very closely controlled with low heat input, small weld beads and interpass temperature controlled to less than 150°C.

Aerospace alloys such as AMS 5858, equivalent to A286, have been produced with improved weldability. The 17/10P grade is particularly sensitive and cannot be welded with matching fillers; a type 312 (29Cr/9Ni) filler gives the best chance of success, although hot cracking in the HAZ may still occur.

Due to the presence of aluminium and/or titanium in many alloys only the inert gas shielded arc welding processes should be used. Some matching composition filler metals are available, again in aerospace grades such as AMS 5804 and these can be aged to give strengths close to those of the parent metal. Alternatively either austenitic, duplex or nickel based weld filler metals may be used.

As is apparent, the metallurgy of these steels can be complex and if there is any doubt concerning welding or heat treatment the advice of specialists should be sought.

Aluminium alloys

Weldability of materials

Job Knowledge

Aluminium and its alloys are used in fabrications because of their low weight, good corrosion resistance and weldability. Although normally low strength, some of the more complex alloys can have mechanical

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properties equivalent to steels. The various types of aluminium alloy are identified and guidance is given on fabricating components without impairing corrosion and mechanical properties of the material or introducing imperfections into the weld.

Material types

As pure aluminium is relatively soft, small amounts of alloying elements are added to produce a range of mechanical properties. The alloys are grouped according to the principal alloying elements, specific commercial alloys have a four-digit designation according to the international specifications for wrought alloys or the ISO alpha - numeric system.

The alloys can be further classified according to the means by which the alloying elements develop mechanical properties, non-heat-treatable or heat-treatable alloys.

Non-heat-treatable alloys

Material strength depends on the effect of work hardening and solid solution hardening of alloy elements such as magnesium, and manganese; the alloying elements are mainly found in the 1xxx, 3xxx and 5xxx series of alloys. When welded, these alloys may lose the effects of work hardening which results in softening of the HAZ adjacent to the weld.

Heat-treatable alloys

Material hardness and strength depend on alloy composition and heat treatment (solution heat treatment and quenching followed by either natural or artificial ageing produces a fine dispersion of the alloying constituents). Principal alloying elements are defined in the 2xxx, 6xxx and 7xxx series. Fusion welding redistributes the hardening constituents in the HAZ which locally reduces material strength.

Processes

Most of the wrought grades in the 1xxx, 3xxx, 5xxx, 6xxx and medium strength 7xxx (e.g. 7020) series can be fusion welded using TIG, MIG and oxyfuel processes. The 5xxx series alloys, in particular, have excellent weldability. High strength alloys (e.g. 7010 and 7050) and most of the 2xxx series are not recommended for fusion welding because they are prone to liquation and solidification cracking.

The technique of Friction Stir Welding is particularly suited to aluminium alloys. It is capable of producing sound welds in many alloys, including those heat treatable alloys which are prone to hot cracking during fusion welding.

Filler alloys

Filler metal composition is determined by:

weldability of the parent metal

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minimum mechanical properties of the weld metal corrosion resistance anodic coating requirementsNominally matching filler metals are often employed for non-heat-treatable alloys. However, for alloy-lean materials and heat-treatable alloys, non-matching fillers are used to prevent solidification cracking.

The choice of filler metal composition for the various weldable alloys is specified in BS EN 1011 Pt 4:2000 for TIG and MIG welding; recommended filler metal compositions for the more commonly used alloys are given in the Table.

Alloy Designation Chemical Designation Classification Filler Application

EN AW-1080A EN AW-Al 99.8(A) NHT R-1080A Chemical plant

EN AW-3103 EN AW-Al Mn1 NHT R-3103 Buildings, heat exchangers

EN AW-4043A EN AW-Al Si5(A) - - Filler wire/rod

EN AW-5083 EN AW-Al Mg4.5Mn0.7 NHT R-5556A Ships, rail wagons, bridges

EN AW-5251 EN AW-Al Mg2Mn0.3 NHT R-5356 Road vehicles, marine

EN AW-5356 EN AW-Al Mg5Cr(A) - - Filler wire/rod

EN AW-5556A EN AW-Al Mg5Mn - - Filer wire/rod

EN AW-6061 EN AW-Al Mg1SiCu HTR-4043AR-5356

Structural, pipes

EN AW-7020 EN AW-Al Zn4.5Mg1 HT R-5556A Structural, transport

HT = Heat treatable, NHT = Non Heat treatable

Imperfections in welds

Aluminium and its alloys can be readily welded providing appropriate precautions are taken. The most likely imperfections in fusion welds are:

porosity  cracking poor weld bead profilePorosity

Porosity is often regarded as an inherent feature of MIG welds; typical appearance of finely distributed porosity in a TIG weld is shown in the photograph. The main cause of porosity is absorption of hydrogen in

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the weld pool which forms discrete pores in the solidifying weld metal. The most common sources of hydrogen are hydrocarbons and moisture from contaminants on the parent material and filler wire surfaces, and water vapour from the shielding gas atmosphere. Even trace levels of hydrogen may exceed the threshold concentration required to nucleate bubbles in the weld pool, aluminium being one of the metals most susceptible to porosity.

To minimise the risk, rigorous cleaning of material surface and filler wire should be carried out. Three cleaning techniques are suitable; mechanical cleaning, solvent degreasing and chemical etch cleaning.

In gas shielded welding, air entrainment should be avoided by making sure there is an efficient gas shield and the arc is protected from draughts. Precautions should also be taken to avoid water vapour pickup from gas lines and welding equipment; it is recommended that the welding system is purged for about an hour before use.

Mechanical cleaning

Wire brushing (stainless steel bristles), scraping or filing can be used to remove surface oxide and contaminants. Degreasing should be carried out before mechanical cleaning.

Solvents

Dipping, spraying or wiping with organic solvents can be used to remove grease, oil, dirt and loose particles.

Chemical etching

A solution of 5% sodium hydroxide can be used for batch cleaning but this should be followed by rinsing in HNO3 and water to remove reaction products on the surface.

Solidification cracks

Cracking occurs in aluminium alloys because of high stresses generated across the weld due to the high thermal expansion (twice that of steel) and the substantial contraction on solidification - typically 5 % more than in equivalent steel welds.

Solidification cracks form in the centre of the weld, usually extending along the centreline during solidification. Solidification cracks also occur in the weld crater at the end of the welding operation. The main causes of solidification cracks are as follows:

incorrect filler wire/parent metal combination incorrect weld geometry welding under high restraint conditionsThe cracking risk can be reduced by using a non-matching, crack-resistant filler (usually from the 4xxx and 5xxx series alloys). The disadvantage is that the resulting weld metal may have a lower strength than the parent metal and not respond to a subsequent heat treatment. The weld bead must be thick enough to

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withstand contraction stresses. Also, the degree of restraint on the weld can be minimised by using correct edge preparation, accurate joint set up and correct weld sequence.

Liquation cracking

Liquation cracking occurs in the HAZ, when low melting point films are formed at the grain boundaries. These cannot withstand the contraction stresses generated when the weld metal solidifies and cools. Heat treatable alloys, particularly 6xxx and 7xxx series alloys, are more susceptible to this type of cracking.

The risk can be reduced by using a filler metal with a lower melting temperature than the parent metal, for example the 6xxx series alloys are welded with a 4xxx filler metal. However, 4xxx filler metal should not be used to weld high magnesium alloys (such as 5083) as excessive magnesium-silicide may form at the fusion boundary decreasing ductility and increasing crack sensitivity.

Poor weld bead profile

Incorrect welding parameter settings or poor welder technique can introduce weld profile imperfections such as lack of fusion, lack of penetration and undercut. The high thermal conductivity of aluminium and the rapidly solidifying weld pool make these alloys particularly susceptible to profile imperfections.

Nickel and nickel alloys

Weldability of materials

Job Knowledge

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Nickel and nickel alloys are chosen because of their:

corrosion resistance heat resistance and high temperature properties low temperature propertiesTypes of nickel alloys are identified and guidance is given on welding processes and techniques which can be used in fabricating nickel alloy components without impairing their corrosion or mechanical properties or introducing flaws into the weld.

Material types

The alloys can be grouped according to the principal alloying elements. Although there are National and International designations for the alloys, tradenames such as Inconel and Hastelloy, are more commonly used.

In terms of their weldability, these alloys can be classified according to the means by which the alloying elements develop the mechanical properties, namely solid solution alloysand precipitation hardened alloys. A distinguishing feature of precipitation hardened alloys is that mechanical properties are developed by heat treatment (solution treatment plus ageing) to produce a fine distribution of particles in a nickel-rich matrix.

Solid solution alloys

Solid solution alloys are pure nickel, Ni-Cu alloys and the simpler Fe-Ni-Cr alloys. These alloys are readily fusion welded, normally in the annealed condition. As the heat affected zone (HAZ) does not harden, heat treatment is not usually required after welding.

Precipitation hardening alloys

Precipitation hardening alloys include Ni-Cu-Al-Ti, Ni-Cr-Al-Ti and Ni-Cr-Fe-Nb-Al-Ti. These alloys may susceptible to post-weld heat treatment cracking.

Weldability

Most nickel alloys can be fusion welded using gas shielded processes like TIG or MIG. Of the flux processes, MMA is frequently used but the SAW process is restricted to solid solution alloys and is less widely used.

Solid solution alloys are normally welded in the annealed condition and precipitation hardened alloys in the solution treated condition. Preheating is not necessary unless there is a risk of porosity from moisture condensation. It is recommended that material containing residual stresses be solution-treated before welding to relieve the stresses.

Post-weld heat treatment is not usually needed to restore corrosion resistance but thermal treatment may be required for precipitation hardening or stress relieving purposes to avoid stress corrosion cracking.

Filler alloys

Filler composition normally matches the parent metal. However, most fillers contain a small mount of titanium, aluminium and/or niobium to help minimise the risk of porosity and cracking.

Filler metals for gas shielded processes are covered in BS EN 18274:2004 and in the USA by AWS A5.14. Recommended fillers for selected alloys are given in the table.

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Table 1: Filler selection for nickel alloys

Parent Alloy Filler designations Comments

AlloyBS EN

ISO 18274

AWS A5.14Trade names

 

Pure nickel        

Nickel 200 Ni 2061 ERNi-1 Nickel 61 Matching filler metal normally contains 3%Ti

Nickel Copper        

Alloy 400 Ni 4060 ERNiCu-7 Monel 60Matching filler metal contains additions of Mn, Ti and Al

Nickel Chromium        

Brightray S Ni 6076 - NC 80/20Ni-Cr and Ni-Cr-Fe filler metals may be used

Nimonic 75 Ni 6076 - NC 80/20

Nickel-Chromium-Iron

       

Alloy 800 Ni 6625ERNiCrMo-3

Inconel 625Thermanit 21/33

Usually welded with Ni-Cr-X alloys, but more nearly matching consumables are available which contain higher C and also Nb

Alloy 600 Ni 6082 ERNiCr-3 Inconel 82 Matching filler metal contains Nb addition

Alloy 718 Ni 7718 ERNiFeCr-2 Inconel 718Matching filler metal is normally used but Alloy 625 is an alternative consumable , if postweld heat treatment is not applied

Nickel-Chromium-Molybdenum

       

Alloy 625 Ni 6625ERNiCrMo-3

Inconel 625Filler metal is also used widely for cladding and dissimilar welds

Hastelloy C-22 Ni 6022ERNiCrMo-10

Hastelloy C-22

 

Nickel-Molybdenum

       

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Hastelloy B-2 Ni 1066 ERNiMo-7Hastelloy B-2

Corrosion resistant alloys require matching fillers

Imperfections and degradation

Nickel and its alloys are readily welded but it is essential that the surface is cleaned immediately before welding. The normal method of cleaning is to degrease the surface, remove all surface oxide by machining, grinding or scratch brushing and finally degrease.

Common imperfections found on welding are:

porosity oxide inclusions and lack of inter-run fusion weld metal solidification cracking microfissuringAdditionally, precautions should be taken against post-welding imperfections such as:

post-weld heat treatment cracking stress corrosion crackingPorosity

Porosity can be caused by oxygen and nitrogen from air entrainment and surface oxide or by hydrogen from surface contamination. Careful cleaning of component surfaces and using a filler material containing deoxidants (aluminium and titanium) will reduce the risk.

When using argon in TIG and MIG welding, attention must be paid to shielding efficiency of the weld pool including the use of a gas backing system. In TIG welding, argon-hydrogen gas mixtures tend to produce cleaner welds.

Oxide inclusions and lack of inter-run fusion

As the oxide on the surface of nickel alloys has a much higher melting temperature than the base metal, it may remain solid during welding. Oxide trapped in the weld pool will form inclusions. In multi-run welds, oxide or slag on the surface of the weld bead will not be consumed in the subsequent run and may cause lack of fusion imperfections.

Before welding, surface oxide, particularly if it has been formed at a high temperature, must be removed by machining or abrasive grinding; it is not sufficient to wire brush the surface as this serve only to polish the oxide. During multipass welding, surface oxide and slag must be removed between runs.

Weld metal solidification cracking

Factors which control solidification cracking include alloy, welding process and welding conditions. For example, solidification cracking is a factor which limits the application of submerged arc welding, both with respect to applicable alloys and welding conditions. More generally, this type of cracking leads to restriction of weld shape, welding speed and technique.

Microfissuring

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Similar to austenitic stainless steel, nickel alloys are susceptible to formation of liquation cracks in reheated weld metal regions or parent metal HAZ. This type of cracking is controlled by factors outside the control of the welder such as grain size or impurity content. Some alloys are more sensitive than others. For example, some cast superalloys are difficult to weld without inducing liquation cracks.

Post-weld heat treatment cracking

This is also known as strain-age or reheat cracking. It is likely to occur during post-weld ageing of precipitation hardening alloys but can be minimised by pre-weld heat treatment. Solution annealing is commonly used but overageing gives the most resistant condition. Alloy 718 alloy was specifically developed to be resistant to this type of cracking.

Stress corrosion cracking

Welding does not normally make most nickel alloys susceptible to weld metal or HAZ corrosion. However, when Alloy 400 will be in contact with caustic soda, fluosilicates or HF acid, stress corrosion cracking is possible. For such service, thermal stress relief is applied after welding.

Stress corrosion can also occur in Ni-Cr alloys in high temperature water. High chromium filler metal has been developed for welds and overlays in this environment.

Welding of nickel alloys - Part 1

Job Knowledge

Nickel is a relatively simple metal. It is face centred cubic and undergoes no phase changes as it cools from melting point to room temperature; similar to a stainless steel. Nickel and its alloys cannot therefore be hardened by quenching so cooling rates are less important than with, say, carbon steel and preheating if the ambient temperature is above 5°C is rarely required. Nickel and its alloys are used in a very wide range of applications - from high temperature oxidation and creep resistance service to aggressive corrosive environments and very low temperature cryogenic applications. Nickel may be used in a commercially pure form but is more often combined with other elements to produce two families of alloys - solid solution strengthened alloys and precipitation hardened alloys. Typical compositions of some of the more common alloys are given in the Table.

Table. Typical composition and properties of some of the more common alloys

Alloy designation Alloy type Typical chemical composition % Mechanical properties

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    Ni Cr Mo Fe Nb Al Ti Others 0.2% proof, MPa UTS, MPa El, %Alloy 200 CP 99.2 - - 0.2 - - - Mn 0.3 148 452 45Monel® 400 SS 68 - - 1.75 - - - Cu 33 235 562 38Monel® K500 PH 65 - - 1.25 - 2.950.55Cu 32 795 1100 18Alloy 600 SS 75 15.5 - 8.5 - - - - 305 670 40

Alloy 617 SS 46 22 9 0.75 - 1.250.45Co 12.5B 0.004 345 725 60

Alloy 625 SS 64 22 8 2.75 3.650.250.25 - 472 920 45Alloy 718 PH 52 19 3 Rem 5.2 0.5 0.95 - 1100 1420 18Alloy 800 SS 32 22 - 42 - 0.450.45 - 290 605 42Alloy 825 SS 42 21.53 28 - 0.1 0.9 Cu 2.25 330 715 39Alloy C276 SS 55 15.516 5.5 - - - W 3.75 345 795 60

Nimonic® PE16 PH 44 16.73.3 29 - 1.2 1.2 B 0.004Zr 0.03 450 825 28

All the conventional welding processes can be used to weld nickel and its alloys and matching welding consumables are available. As mentioned above, nickel and its alloys are similar in many respects to the austenitic stainless steels; welding procedures are likewise also similar. Nickel, however, has a coefficient of thermal expansion less than that of stainless steel so distortion and distortion control measures are similar to those of carbon steel.

The most serious cracking problem with nickel alloys is hot cracking in either the weld metal or close to the fusion line in the HAZ with the latter being the more frequent. The main source of this problem is sulphur but phosphorus, lead, bismuth and boron also contribute. Both weld metal and HAZ cracking are generally the result of contamination by grease, oil, dirt, etc left behind following inadequate cleaning; excess sulphur in the parent or weld filler metals causing a problem is a rare event. Machining or vigorous stainless steel wire brushing followed by thorough degreasing with a suitable solvent is necessary prior to welding, with the welding taking place within about eight hours to reduce the risk of contamination. Any heat treatment must be carried out using sulphur-free fuel or by using electric furnaces. Components that have been in service and require weld repair may need to be ground or machined prior to degreasing to remove any contaminants that have become embedded in the surface in or adjacent to the weld repair area. Remember that if mechanical wire brushing is carried out AFTER the degreasing operation or during welding the compressed air from air powered tools contains both moisture and oil and the cleaned surfaces may be therefore be re-contaminated.

Porosity can be a problem with the nickel alloys, the main culprit being nitrogen. As little as 0.025% nitrogen will form pores in the solidifying weld metal. Quite light draughts are capable of disrupting the gas

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shield and atmospheric contamination will occur resulting in porosity. Care must be taken to ensure that the weld area is sufficiently protected and this is particularly relevant in site welding applications. With the gas shielded processes, gas purity and the efficiency of the gas shield must be as good as possible. Gas hoses should be checked for damage and leaks at regular intervals and, with the TIG process, as large a ceramic shroud as possible should be used together with a gas lens. It goes without saying that gas purging of the root is essential when depositing a TIG root pass.

A small amount of hydrogen (up to 10%) added to the argon shield gas has been found to reduce the problem. Start and finish porosity is a problem when MMA welding. The weld start should be carried out by welding back over the arc strike position, remelting any porosity that has formed due to the poor gas shielding at the start of the weld. Care also needs to be taken at the weld end, with the arc length reduced and travel speed increased slightly to reduce weld pool size.

Oxygen is also a cause of porosity in certain circumstances when it combines with carbon in the weld pool to form carbon monoxide. Consumable manufacturers generally overcome this problem by ensuring that sufficient deoxidants (primarily manganese, aluminium and titanium) are present in the filler metal.

One feature of nickel alloys that is often encountered is the formation on the surface of the weld pool of a viscous and adherent scum. This can be difficult to remove and can result in inclusions and lack of inter-run fusion if not removed prior to depositing the next pass. Wire brushing is frequently not sufficient to remove this layer and it then becomes necessary to grind the weld surface.

The weld pool, in addition to this surface film, is also sluggish and does not flow freely as with a carbon or stainless steel. This may result in a lumpy and very convex weld bead and a poor toe blend unless the welder manipulates the weld pool to avoid such defects. Although stringer beads may be used, a slight weave to assist the weld metal to wet the side walls of the preparation is beneficial. In addition, weld preparations must be sufficiently wide to enable the welder to control and direct the weld pool; an included angle of 70 to 80° is recommended for V butt welds.

A U preparation included angle of 30 to 40° is acceptable and, though more expensive to machine than a V preparation, may be cheaper overall as the amount of filler wire required can be reduced, depending on material thickness. Addition of hydrogen to the shield gas (up to 10%H in argon) in TIG welding also has been found to be beneficial in reducing the weld pool surface tension.

A further characteristic of nickel alloys is that the amount of penetration is less than with a carbon or stainless steel. Increasing the welding current will not increase penetration. The implication of this is that the root face thickness in single sided full penetration welds should be less than with a stainless steel. It is recommended that the thickness of the root face should not be greater than 1.5mm in a zero gap TIG butt weld. Removable backing strips are very useful to control root bead shape. These can be made from copper, stainless steel or a nickel alloy. Carbon or low alloy steel backing strips should be avoided.

Although weldability of nickel and its alloys is generally good the composition, metallurgical structure and its heat treatment and/or service history all affect its response to welding. Wrought, fine grained components have better weldability than cast items as these often have significant amounts of segregation. Coarse grains may lead to micro-fissuring in the HAZ thus high heat input is best avoided. All the alloys are best welded in the annealed or solution treated condition and this applies particularly to the precipitation hardenable alloys such as Inconel 718.

Welding of nickel alloys - Part 2

Job Knowledge

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In Part 1 the importance of cleanliness, particularly the removal of all sulphur containing compounds, was mentioned. With respect to defect free welding of nickel and its alloys this cannot be over-emphasised. 

As well as sulphur, however, there are several other substances that can lead to embrittlement of the nickel alloys when they are exposed to high temperatures. Amongst these are lead, phosphorus, boron and bismuth.

These may be present in oils, grease, cutting fluids, paints, marker pen inks, temperature indicating crayons, etc; it may not be possible to avoid using these during fabrication so it is essential that these are removed if the component is to be welded, heat treated or is to enter high temperature service.

Fuel gases frequently contain sulphur and it may be necessary to use radiant gas heaters or electrical elements for local heating or in heat treatment furnaces.

Nickel alloys can be welded using all the conventional arc welding and power beam processes, the commonest processes being TIG or MIG with pure argon, argon/hydrogen or argon/helium mixtures as shield gases and MMA where basic flux coatings provide the best properties.

However, if argon/helium mixes are used it is only when there is more than 40% helium that any significant benefits with respect to penetration and improved fusion will be noticed. Submerged arc welding is restricted to welding solid solution alloys using basic fluxes. Matching welding consumables are available for most of the nickel alloys. See Job Knowledge 22 for recommendations for a range of alloys.

Slag from MMA welding and particularly submerged arc welding can be difficult to remove from the nickel alloys and often needs to be ground between runs to remove it completely. It is also often necessary to grind the surface of each run when welding with the gas shielded processes to remove oxide scabbing, wire brushing simply polishing these oxides.

Failure to remove slag or oxide scabs will result not only in weld metal inclusions but also reduce corrosion resistance if left on exposed surfaces. Total welding times can therefore be substantially longer than the equivalent joint in stainless or carbon steel and welders need to be fully acquainted with these differences when converting from welding steels to nickel alloys.

Comments regarding the recommended weld preparations were included in Part 1. Although the weld preparations are similar to those used for steel it is worth considering the use of double V or U type preparations at thicknesses less than would be considered with steels. The additional cost of the preparation is offset by savings in consumable costs (nickel being an expensive metal) and welding time.

The majority of nickel alloys are best welded in the annealed or solution treated condition, particularly if the alloys have been cold worked. As mentioned in Part 1, preheat is not required except to remove condensation or if the ambient temperature is below about 5°C when a moderate preheat of 40-50°C is recommended.

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Interpass temperature should not be allowed to rise above 250°C although some alloy suppliers recommend an interpass as low as 100°C for certain alloys such as Alloy C276.

Remember the potential hot crack problems if thermal crayons are used to measure this temperature! For most alloys heat input should be controlled to moderate levels (say 2kJ/mm maximum) to limit grain growth and HAZ size although for some Alloys 718, C22, and C276 for example, a maximum heat input of 1kJ/mm is recommended.

Conversely if too fast a travel speed is used in an attempt to maintain a low heat input this can result in a narrow weld bead sensitive to centre line cracking. Adequate testing during welding procedure development should be used to optimise the range of acceptable welding parameters.

The solid solution alloys such as Alloy 200 or 625 do not require post weld heat treatment to maintain corrosion resistance but may be subject to PWHT either to reduce the risk of stress corrosion cracking if the alloy is to be used in caustic soda service or in contact with fluoro-silicates or to provide dimensional stability.

A typical stress relief treatment would be 700°C for ½ an hour for Alloy 200; 790°C for four hours for the higher chromium content alloys such as Alloy 600 or 625.

The nickel-molybdenum alloys are identified with the prefix B eg B1, B2, etc. and are used in reducing environments, such as hydrogen chloride gas and sulphuric, acetic and phosphoric acids. Alloy B2 is the most frequently encountered alloy and matching filler metals are available. Unlike Alloy B1, Alloy B2 does not form grain boundary carbide precipitates in the weld heat affected zone, so it may be used in most applications in the as-welded condition.

Alloy 400, a 70Ni-30Cu alloy, has good corrosion resistance when exposed to hydrofluoric acid, strong alkaline solutions and sea water.

A matching filler metal, Alloy 190, is available but this can become anodic in salt solutions, leading to galvanic corrosion and it is recommended that one of the Ni-Cr alloy fillers such as Alloy 600 or 625 is used in this environment.

The age hardened alloy K-500 does not have a matching filler metal and is generally welded using the Alloy 190 filler, the reduction in strength being taken into account during the design phase.

Precipitation hardened alloys are best welded in the solution treated condition; welding these alloys in the age hardened condition is likely to result in HAZ cracking.

The ageing process in the alloys is sufficiently sluggish that the components can be welded in the solution treated condition and then aged at around 750°C without the mechanical properties being degraded.

A solution treatment of the welded item followed by ageing will provide the highest tensile strength.

The sensitivity of the age hardened alloy to cracking causes problems when attempts are made to repair items, particularly when these have been in high temperature service and additional precipitation on the grain boundaries has occurred.

Little can be done to overcome this problem apart from a full solution heat treatment but this is often not possible with a fully fabricated component. If repair is to be attempted, small weld beads and controlled low heat input welds are recommended.

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If the design permits, a low strength filler metal, eg Alloy 200 or 600, may be used to reduce the risk. Buttering the faces of the repair weld preparation, sometimes combined with a peening operation, has been successful.

Many of the nickel alloy filler metals have been used for making dissimilar metal joints with excellent results; dilution when welding joints between ferritic, stainless and duplex steels being less important than when using a type 309 stainless steel filler.

Nickel also has a coefficient of thermal expansion between that of ferritic and austenitic steels and therefore suffers less from thermal fatigue when high temperature plant is thermally cycled. Alloy 625 has been a popular choice, the weld tensile strength matching or exceeding that of the parent metal.

There are limitations to this approach, and caution needs to be exercised when selecting a suitable filler.

For example, Alloy 625 has been extensively used for welding dissimilar joints in austenitic and duplex steels.

Use of this filler metal has resulted in the formation of niobium rich precipitates adjacent to the fusion line and has been discontinued. Alloy 59 or C22 filler metals has replaced Alloy 625 as the filler of choice.

Copper and copper alloys

Weldability of materials

Job Knowledge

Copper and copper alloys are chosen because of their corrosion resistance and electrical and thermal conductivity.

The various types of copper alloys are identified and guidance is given on processes and techniques which can be used in fabricating copper alloy components with a view to maintaining their corrosion or mechanical properties whilst avoiding the introduction of defects into the welds.

Alloy types

The main categories of copper and copper alloy are listed below:

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Table 1. Frequently used copper alloys and recommended filler metals

Alloy type Recommended filler

Coppers (tough pitch, phosphorus deoxidised) Cu 1897, Cu 1898

Brasses (low Zinc) Cu 6328, Cu 6560

Nickel Silvers (20%Zn/15%Ni type) Cu 6328, Cu 6560

Silicon Bronze (3%Si) Cu 6560

Phosphor Bronze (4.5% to 6%Sn/0.4%P) Cu 5180

Aluminium Bronze (<7.8%Al) Cu 6240, Cu 6100

Aluminium Bronze (>7.8%Al) Cu 6180, Cu 6328

Aluminium Bronze (6%Al/2%Si) Cu 6100

Gunmetal (low lead) Cu 5180, Cu 6560, Cu 6180

Cupro-Nickel (10%Ni) Cu 7061, Cu 7158

Cupro-Nickel (30%Ni) Cu 7158

Pure copper Copper with small alloy additions (less than 5% in total) Brasses e.g. copper-zinc (Cu-Zn) Nickel silvers e.g. copper-zinc-nickel (Cu-Zn-Ni) Bronzes e.g. copper-tin (Cu-Sn) (phosphor bronze alloys also contain phosphorus) Gunmetals e.g. copper-tin-zinc (Cu-Sn-Zn) (some alloys may contain lead) Aluminium bronze e.g. copper-aluminium (Cu-Al) (most alloys also contain iron and many nickel) Cupro-nickels e.g. copper-nickel (Cu-Ni)The most frequently used copper alloys are listed in Table 1, together with a range of welding electrodes for fusion welding as per BS EN 14640:2005. Similar filler wire compositions are given in AWS A5.7/A5.7M:2008 and covered electrodes are specified in A5.6/A5.6M:2007.

It should be mentioned that welding of Nickel Silvers (45%Zn/10%Ni), leaded Gunmetal and high Zinc Brasses (40%Zn) is not recommended.

Copper alloys have quite different welding characteristics due to differences in thermal conductivity. For example copper, due to its high thermal conductivity, may require substantial preheat to counteract the very high heat sink. However, some of the alloys which have a thermal conductivity similar to low carbon steel, such as cupro-nickel alloys, can normally be fusion welded without a preheat.

Copper

Copper is normally supplied in the form of

oxygen bearing, tough pitch copper

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phosphorus deoxidised copper oxygen-free copperTough pitch copper contains stringers of copper oxide (<0.1% oxygen as Cu2O) which does not impair the mechanical properties of wrought material and it has high electrical conductivity.Oxygen-free and phosphorus deoxidised copper are more easily welded.

TIG and MIG are the preferred welding processes but oxyacetylene and MMA welding can be also used in the repair of tough pitch copper components. Helium and nitrogen-based shielding gases, which have higher arc voltages, can be used as an alternative to argon to counteract the high thermal conductivity of coppers.

Avoiding weld imperfections 

During fusion welding of tough pitch copper, the high oxygen content of the alloy often leads to embrittlement in the heat affected zone (HAZ) and weld metal porosity. Phosphorus deoxidised copper is more weldable but residual oxygen can result in porosity in autogenous welds especially in the presence of hydrogen. Porosity can be avoided by using appropriate filler wire containing deoxidants (Al, Mn, Si, P and Ti).

Thin section material can be welded without preheat. However, over 5mm thickness all grades need preheat to produce a fluid weld pool and avoid fusion defects. Thick section components may need a preheat temperature as high as 600 deg.C.

Copper with small alloying additions

Low amounts of sulphur or tellurium can be added to improve machining. However, these grades are normally considered to be unweldable.

Precipitation hardened alloys contain small additions of chromium, zirconium or beryllium. and have superior mechanical properties. Chromium and beryllium coppers may suffer from HAZ cracking unless they are heat treated before welding. When welding beryllium copper, care should be taken to avoid inhaling the welding fumes, which are poisonous.

Brasses (copper-zinc alloys) and nickel silvers

When considering weldability, brasses can be separated into two groups viz. low zinc (up to 20% Zn) and high zinc (30 to 40% Zn). Nickel silvers contain 20 to 45% zinc and nickel to improve strength. The main problem in fusion welding these alloys is the volatilisation of the zinc which results in white fumes of zinc oxide and weld metal porosity. Only low zinc brasses are weldable using fusion welding processes such as TIG and MIG.

Avoiding weld imperfections

To minimise porosity, a zinc-free filler wire should be used, either silicon bronze (Cu 6560) or an aluminium bronze (Cu 6180). High welding speeds will reduce pore size.

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TIG and MIG processes are used with argon or an argon-helium mixture but not with nitrogen. Preheat is normally used for low zinc (<20% Zn) to avoid fusion defects due to the high thermal conductivity,. Although preheat is not needed for higher zinc content alloys, slow cooling reduces cracking risk. Post weld heat treatment also helps to reduce the risk of stress corrosion cracking in areas where restraint is high.

Bronzes (tin bronze, phosphor bronze, silicon bronze and gunmetal)

Tin bronzes typically contain between 1% to 10% tin. Phosphor bronze contains up to 0.4% phosphorus. Gunmetal is essentially a tin bronze with up to 5% zinc and it may have lead additions up to 5%. Silicon bronze contains 3% silicon and 1% manganese approximately and it is probably the easiest of the bronzes to weld.

Avoiding weld imperfections

Matching filler compositions are normally employed for welding bronzes. Autogenous welding of phosphor bronzes is not recommended due to weld metal porosity. However, this risk can be reduced by using a filler wire with a higher level of deoxidants. Gunmetal is not considered weldable since it is susceptible to hot cracking.

Aluminium bronze

There are essentially two types of aluminium bronzes; single phase alloys containing between 5 to 10% aluminium, with a small amount of iron or nickel, and more complex, two phase alloys containing up to 12% aluminium and about 5%of iron with specific alloys also containing nickel, manganese and silicon. Gas shielded welding processes are preferred for welding this group of alloys. In TIG welding, the presence of a tenacious, refractory oxide film requires AC(argon), or DC with a helium shielding gas. Due to its low thermal conductivity, a preheat is not normally required except when welding thick section components.

Avoiding weld imperfections

Rigorous cleaning of the material surface is essential, both before and after deposition of each welding pass, to avoid porosity. Single phase alloys can be susceptible to weld metal and HAZ cracking under highly restrained conditions. It is often necessary to use matching filler metals to maintain corrosion resistance but a non-matching, two phase, filler can also reduce the risk of cracking. Two phase alloys are easier to weld. For both types, preheat and interpass temperatures should be controlled carefully to prevent cracking.

Cupro-nickels

Cupro-nickel alloys contain 5 to 30% nickel with specific alloys having additions of iron and manganese; 90/10 and 70/30 (Cu/Ni) alloys are commonly welded grades. These alloys are single phase and generally considered to be weldable using inert gas processes and, to a lesser extent, MMA. A matching filler is normally used. 70/30 (Cu 7158) is often regarded as a 'universal' filler for these alloys. The thermal conductivity of cupro-nickel alloys is similarto low carbon steels, and therefore preheating is not required.

Avoiding weld imperfections

Cupro-nickels do not contain deoxidants, and therefore, autogenous welding is not recommended due to the risk of porosity. Filler metal compositions contain typically 0.2 to 0.5% titanium, to minimise weld metal porosity. Argon shielding gas is normally used for both TIG and MIG but in TIG welding, an argon-hydrogen mixture, with appropriate filler, improves weld pool fluidity and produces a cleaner weld bead. Gas backing (usually argon) is recommended, especially in pipe welding, to produce an oxide-free underbead.

Welding of copper and its alloys - Part 1

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Job Knowledge

Repair of copper boiler from the Flying Scotsman 

Of all metals copper is the most ancient, having been first used to fabricate tools and weapons since about 3500 years BC. Welders and metallurgists can therefore claim to have a very long pedigree! Pure copper is soft, ductile and easily worked but can be strengthened only by cold working. It does not undergo phase changes so cannot be hardened by heat treatment as can a steel. This also applies to many of the copper alloys so that any application of heat will soften the cold worked alloy, resulting in a significant loss of strength in the heat affected zones.

Two additional characteristics of copper and some of its alloys are

1. high thermal conductivity, meaning that preheat is required for many joints, even at quite modest thicknesses, and

2. the high coefficient of thermal expansion, meaning that distortion can be an issue with root gaps rapidly closing during welding. 

Alloying with a range of metals can be used to improve the mechanical properties and/or corrosion resistance. These alloys can be conveniently placed into nine separate groups as listed below. In addition to those listed there are several grades of free machining alloys containing lead (Pb) or selenium (Se). These free machining grades are hot-short and very sensitive to hot cracking. They are best avoided by the welder although they can be successfully joined by brazing or soldering.

Pure copper with less than 0.7% residual elements High copper alloys with less than 5% alloying elements Copper alloys with up to 40% zinc (Zn) (brasses) Copper alloys with less than 10% tin (Sn) (bronzes) Copper alloys with less than 10% aluminium (Al) (aluminium bronzes often shortened to ally-

bronze) Copper alloys with less than 3% silicon (Si) (silicon bronze) Copper alloys with less than 30% nickel (Ni) (cupro-nickel alloys) Copper alloys with less than 40%Zn and less than 18%Ni (nickel silvers) Copper alloys with less than 10%Sn and less than 4%Zn (red brass or gunmetal) Special alloys containing1. 0.1-1.5% cadmium (Cd)

2. less than 2.7% beryllium (Be)

3. 0.6-1.2% chromium (Cr)

4. 0.1-0.2% zirconium (Zr).

This group of special alloys are capable of being precipitation hardened.

Copper alloys can be welded with most of the conventional welding processes although of the arc welding processes, gas shielded arc methods are the most common.

Pure copper alloys

There are three separate grades of pure copper: Oxygen-free copper with less than 0.02% oxygen; tough pitch copper that contains <0.1% of oxygen, present as copper oxide, and phosphorous (P) deoxidised copper with 0.05% P up to 0.05% arsenic (As). Oxygen-free copper has the highest electrical conductivity,

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P-deoxidised copper is the alloy most frequently used for pressure vessel and heat exchangers. Oxygen-free copper is the most readily weldable although porosity may be a problem if non-deoxidised filler metals are used.

The copper oxides in tough pitch copper can result in embrittlement of the heat affected zones due to oxide films forming on the grain boundaries. Weld metal porosity, even when using fully deoxidised filler metals, is also a major problem caused by the dissociation of the copper oxide, particulaly when hydrogen (H) is present.

Phosphorus deoxidised copper presents less of a porosity problem although weld metal porosity is still likely to be formed, particularly in autogenous welds. It is essential therefore that filler metals contain strong deoxidants, the commonest being silicon (Si) and manganese (Mn). Hydrogen control is also necessary so correctly baked low hydrogen electrodes are necessary when manual metal arc welding. Clean, grease-free wires and rods and high purity shield gases are required when TIG or MIG welding.

The two filler metals most often selected to weld the pure copper alloys are AWS A5.7 ERCu, the C7 of the now superceded BS 2901 Part 3 and ERCuSi-A, the old C9 of BS 2901. ERCu typically contains 0.4% of Si and Mn with 0.8% of Sn to aid fluidity; ERCuSi-A contains 1%Mn and 3%Si and is the preferred filler metal for tough pitch and P-deoxidised copper. BS 2901 Part 3 has been replaced by BS EN ISO 24373:2009 Welding consumables. Solid wires and rods for fusion welding of copper and copper alloys.

Shielding gases for welding are argon, helium and nitrogen or mixes of two or more of these. Pure argon may be used for TIG welding up to a thickness of some 2mm and for MIG welding up to approximately 5mm - above these thicknesses an argon-helium mixture will give better results with greater heat input and less risk of lack of fusion defects.

Nitrogen and argon-nitrogen gas mixes have been used in the past with some advantages being gained in terms of increased heat input from the high voltage nitrogen arc but such gases are not commercially available and argon-helium or helium shield gases are now the preferred choice. The high thermal conductivity of copper means that not only are high heat input shielding gases required as thickness increases, but preheat is necessary at section thicknesses exceeding 2mm. A very rough guide to recommended preheat and welding current levels is given in the table for TIG and MIG welding.

Process Thickness (mm) Shielding Gas Preheat °C Welding Current (amps)

TIG        

  1.0 argon >10 20 - 60

  1.0 - 2.0 argon >10 50 - 160

  2.0 - 5.0 argon/75helium 50 120 - 300

  6.0 - 10.0 argon/75helium 100 - 200 250 - 375

  12.5 argon/75helium 350 350 - 420

  15.0 argon/75helium 400 - 450 400 - 470

MIG        

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  <5.0 argon 10 - 100 175 - 240

  5.0 - 7.0 argon/75helium 100 250 - 320

  10.0 - 12.5 argon/75helium 200 - 300 300 - 400

  >16.0 argon/75helium 350 - 450 350 - 600

When welding thick copper with preheats of over 250°C and welding currents of more than 350 amps then the health and safety of the welder and personnel working in the vicinity must be considered.

Lagging the item being welded with thermal blankets is essential as is the provision of adequate screening from the very powerful TIG or MIG arc. The welder should select a dense filter glass of at least shade 13 when using welding currents above 300 amps to reduce eye strain.

Typical butt weld preparations are:-

up to 1.5mm thickness - square edge,no gap 1.5 to 3mm - square edge with 1.5mm gap 3 to 12mm single -V, included angle of 60° to 90°, feather edge and up to a 1.5mm gap 12mm to 25mm single V, included angle of 60 to 90°, 1.5 to 3mm root face, 1.5mm maximum gap Over 25mm thickness double V, included angles of 60 to 90°, 1.5 to 3mm root face, 1.5mm

maximum gapCarbon, stainless steel or ceramic tiles or tape can be used as temporary backing strips and are helpful in controlling root bead shape.

The Job Knowledge series is aimed at the welder and therefore tends to concentrate on the conventional arc welding processes. It is worth bearing in mind that electron beam and friction welding, including friction stir, have been used extensively and very successfully to weld thick section copper without the need for filler metals, high preheat temperatures and expensive shielding gases.

Copper alloys - brasses and bronzes

Job Knowledge

The main alloying element in the brasses is zinc (Zn). There are three families; brass with zinc content less than 20%, high zinc alloys with 30-45% zinc and the nickel-silvers that contain 20-45% zinc and 20% nickel. These alloys are available as wrought or cast products, the low Zn alloys being used generally for jewellery and coins, the higher Zn alloys in applications where increased mechanical strength is required such as plumbing products, pump casings and thin wall low pressure vessels. Nickel silver, as the name suggests, is a less expensive alternative to silver (Ag) and is used for jewellery, coinage and cutlery. On an historical note, the panels of the 1907 Rolls Royce Silver Ghost are made from nickel silver, hence the name.

With the exception of brasses containing lead (Pb) all the brasses are weldable, the low zinc alloys being the easiest. The main problem with welding the alloys is weld metal porosity caused by the zinc boiling off during melting. Zinc melts at 420°C and boils at 910°C so brazing using an oxy-acetylene torch and a copper-silver filler is a possible alternative to welding, being capable of providing joints with adequate mechanical properties and without the porosity problems. Boiling the zinc may also result in large amounts of zinc oxide in the welding fume and this can be a health and safety issue. Brasses may be welded using MMA, MIG or TIG. Filler metals are available although these are generally based on copper-silicon or copper-tin alloys due to the problems of transferring zinc across the welding arc. A typical MIG/TIG filler

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metal would be the 3% silicon alloy specified in EN ISO 24373 SCu 6560 (CuSi3Mn1). Successful welds can also be made using copper-tin alloys such as Cu-7%Sn and Cu-12%Sn. These can be obtained as both MIG/TIG wires and as MMA electrodes.

The Cu-Si filler metal flows easily and a 60° included angle weld preparation should give acceptable results. The Cu-Sn weld metal is more sluggish and an included angle of at least 70° is advisable. The shielding gas used for MIG or TIG welding of thin section components is high purity argon. In thicker sections, over 5mm thick, the addition of helium will greatly assist in providing sufficient heat for full fusion as will the use of pulsed welding current. Brass, like copper, has a high coefficient of thermal conductivity. TIG welding is generally limited to joint thickness of around 10mm, MIG being the preferred process for thicker sections. Preheating to between 100 and 300°C, depending upon section thickness can be helpful in reducing zinc loss, particularly in the high zinc alloys, by enabling a lower welding current to be used, resulting in less melting of the parent metal.

There is a potential problem in service of stress corrosion, known as season cracking, in mildly corrosive media such as ammonia or sea water due to the residual stresses from welding. This can be largely dealt with by annealing the welded item at 260-300°C.

The next group of alloys is the bronzes. These may be alloyed with tin, generally described as phosphor bronze, silicon or aluminium. Many of these alloys, like the brasses, are alloyed with lead to improve machinability. These leaded alloys are generally regarded as unweldable and specialist advice should be sought if the need arises.

Phosphor bronze alloys contain between 1 and 12% tin with a small amount (0.01-0.1%) of phosphorus (P) when this is used solely as a deoxidising agent. True phosphor-bronzes contain at least 0.1%P and as much as 1.0%P in some of the cast phosphor bronzes.

The alloys are corrosion resistant and have excellent wear characteristics so they are used for valves, bearings and machine parts. From a weldability point of view the main problem is that the alloys are sensitive to hot cracking and the lower P content alloys are also prone to form oxide films on the weld pool. High welding heat inputs, high preheat and slow cooling rates should therefore be avoided. MIG and TIG welding are the preferred welding processes with argon or helium-argon mixtures. MIG is more suitable than TIG for welding heavier section joints and positional welding is best achieved using pulsed current. Filler metals matching the composition of the parent metal, e.g. EN ISO 24373 CuSn6P, are available. Although MMA welding consumables are available the process is not widely used. A stringer bead welding technique is generally necessary and heavy sections require preheat and interpass temperatures of around 200°C.

Silicon bronzes are probably the easiest of all the bronzes to weld. They contain between 1.0 and 4.0% silicon with small amounts, less than 1.5% in total, of zinc, manganese and/or iron. They have good strength and excellent corrosion resistant properties and are frequently used for heat exchanger tubing, marine hardware and in chemical process plant applications.

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Unlike many of the other copper alloys thermal conductivity is relatively low and this makes it possible to use high welding speeds and to dispense with preheat for the thicker joints. One undesirable characteristic, however, is that the silicon tends to form an oxide film on the weld pool surface that requires vigorous wire brushing of individual weld passes during multi-run welding. There is also a slight tendency to hot shortness at elevated temperatures. It is advisable to stress relieve or anneal components prior to welding and to cool rapidly through the 1000-850°C temperature range.

As with the other bronzes, MIG or TIG welding are the processes of choice using pure argon as the shield gas and consumables that match the parent metal composition, e.g. EN ISO 24373 CuSi3Mn1. Low thermal conductivity means that helium mixes are not necessary and the TIG process can be used for welding components up to 25mm thickness at welding currents of 300amps. However, it should be noted that the weld pool size should be restricted to provide a fast cooling rate.

The last alloy in this series is aluminium-bronze. This family of alloys have compositions between 3 and 15% of aluminium with additions of iron, manganese and nickel. The alloys with less than 8%Al are single phase; those with more than 9%Al are two phase and capable of being quench hardened to give a martensitic micro-structure. All the alloys have excellent corrosion resistance, particularly in marine environments, and are used for pump bodies, valves, bearings and ships propellers.

The characteristic that gives the alloy its corrosion resistance is the strong tenacious aluminium oxide film that forms on the surface. This causes problems of oxide film entrapment and lack of fusion during welding and must be removed. Scraping and wire brushing the surfaces before welding is necessary. With respect to the welding processes, IG and TIG are preferred. With MIG there is no problem in dispersing the oxide film, the DC+ve current breaking up and dispersing the film. DC-ve TIG welding does not provide this cleaning action and it is necessary to use AC-TIG. Inverter-based square wave TIG power sources will give the best control. Argon is the recommended shield gas although a helium/argon mixture may be useful when welding very thick section joints with the MIG process. MMA welding is possible although the fluxes required to remove the oxide film are very aggressive and may cause corrosion problems if not completely removed before the item enters service.

Aluminium bronzes with less than 8% aluminium are prone to hot cracking at temperatures around 700°C and care needs to be taken to reduce residual stresses as much as possible by ensuring accurate fit-up and minimal root gaps. Low heat input procedures should be used and interpass temperature limited to 150°C. These alloys do not require preheat. A filler metal with around 8 to 10% aluminium such as EN ISO 24373 CuAl10Fe1 or AWS A5.7 CuAl-A2 is the best choice as this composition is relatively resistant to hot cracking.

The two phase alloys, i.e. those with more than approximately 9%Al, have very high tensile strengths although the very highly alloyed suffer from a substantial loss of ductility. All the alloys are, however, readily weldable and relatively insensitive to hot cracking. Heat input control is therefore less important although a maximum interpass temperature of 250°C is recommended and a preheat of 150°C may be used when MIG welding thick section joints. AWS A5.7 ER CuAl-A2 (EN ISO 24373 CuAl10Fe1) or, for higher strength, ER CuAl-A3 (EN ISO 24373 CuAl11Fe3) are readily available MIG/TIG filler metals.

Post weld heat treatment is rarely necessary but can be of benefit if the welded item is to experience very corrosive conditions. In this case a stress relief operation at 300-350°C may be beneficial, although precise temperatures and times will depend upon the specific alloy composition, thickness etc. It is possible for the high aluminium duplex alloys to be quenched from 950°C and tempered at 650°C to restore full corrosion resistance but this is rarely done due to cost and distortion issues.

Copper-nickel alloys

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Job Knowledge

Copper and nickel are completely soluble in each other, giving rise to a range of alloys that includes both copper-nickel (Cu-Ni) and nickel-copper alloys, the latter alloys having been covered in the Job Knowledge articles numbers 107 and 108. 

Although there is a wide range of alloys, only two are commercially significant. These are the 90/10 and 70/30 grades and the Table shows typical compositions and mechanical properties. Both grades have excellent corrosion resistance, particularly in sea water applications, and are used extensively in marine and offshore applications. The 70Cu/30Ni alloy is the stronger of the two with a yield strength in the annealed condition of ~150MPa compared with 120MPa of the lower nickel alloy. The 90Cu/10Ni grade however is probably the most used grade as it is less expensive than the higher nickel alloy.

The alloys are single phase and cannot be hardened by heat treatment. The only method of increasing tensile strength is by cold working which, when the metal is in the fully hard condition, can match that of good quality carbon steel. Work hardening, however, has implications with respect to welding in that there will be some strength loss in the HAZs. Fortunately this region is relatively narrow due to the low coefficient of thermal conductivity; approximately the same as steel. This narrow, low strength region can cause problems during welding procedure qualification testing of transverse bend coupons, most of the deformation being concentrated in the narrow area of strength loss. Bend testing is therefore generally carried out using a longitudinal bend specimen.

The other main alloying elements are manganese, around 1%, that is used as a deoxidant and desulphuriser, and up to 2% iron which is added to improve erosion resistance. Some of this iron, perhaps 1% or more, may be replaced with chromium to increase the strength. Niobium may also be added to castings to increase the strength and at the same time improve weldability

Due to a deficiency in deoxidants in the alloys, porosity is a problem and they cannot be welded autogenously. A highly deoxidised filler metal needs to be used although there is an exception to this rule; thin sheet containing substantial amounts of titanium. A very strong deoxidant, is now available and this is capable of being welded autogenously by TIG, plasma-TIG and the power beam processes without significant porosity problems

There are filler metals available that match both grades but it is generally 70Cu/30Ni filler that is used; AWSA 5.6 ECuNi MMA electrodes and AWS A5.7 ERCuNi for TIG and MIG wires. The weld metal from these filler metals overmatches the strength of both grades in the annealed condition. Having a 0.2% proof of some 270MPa it has better handling characteristics than the 90Cu/10Ni filler and is noble with respect to the 90Cu/10Ni parent metal. The 90Cu/10Ni filler metals have a lower 0.2% proof of around 200MPa and should be used for welding 90Cu/10Ni alloys only.

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The weld metal from both grades of filler metal is more sluggish than, say, carbon steel. Weld preparations therefore need to be more open to enable the welder to control and manipulate the weld pool. An included angle of 70 to 80O is recommended. Root face dimensions would typically be 0-1.5mm root face with a zero-1.5mm root gap.

As mentioned above, porosity when welding either grade can be a problem and to reduce the risk the filler metals contain substantial amounts (around 0.5%) titanium. Cleanliness of weld preparations and filler wires is also important, as is the use of high purity shielding gas. Weld preparations may need to have the tenacious oxide films removed by belt or disc sanding and should be thoroughly degreased with commercially available solvents. Stainless steel wire brushes and stainless steel wire wool are also useful.

This cleaning equipment must not be used on any other metals otherwise cross contamination will occur. Ideally the Cu/Ni fabrication area should also be physically separated from other fabrication areas to prevent dust from activities such as grinding settling on the cleaned weld preparations. One point worth noting is, if air powered tools are used for wire brushing or sanding, these may leave a film of moisture and/or oil on the surface (compressed air is seldom completely free of contaminants) and this may result in porosity and or cracking.

Depositing a pore-free root pass can be particularly difficult. Insufficient filler metal coupled with a large amount of dilution from the parent metal may result in unacceptable porosity. Copious amounts of filler metal and a larger than normal root gap (~2-3mm) will reduce porosity to acceptable levels.

Other causes of porosity may be associated with inadequate gas shielding. When TIG welding, use as large a diameter ceramic as possible, together with a gas lens. Arcs should be kept short; too long an arc length may permit atmospheric contamination.

Both the alloys are sensitive to hot cracking. As with the other nickel alloys the main culprit is sulphur but lead, phosphorus and carbon will also have and adverse effect. Cleanliness, as discussed above, is therefore crucial and all grease, oil, marker crayon, paint etc must be removed from the weld preparation and the adjacent areas before welding. To reduce further the risk of hot fissuring the interpass temperature should be limited to 150OC.

The alloys have high coefficients of thermal expansion and more extensive tack welding than would be required for a carbon steel is necessary to prevent excessive distortion and root gaps closing up during welding. Tacks should be wire brushed or ground to bright metal if they are to be incorporated in the completed weld.

TIG (GTAW) welding will give the best quality weld metal and a well shaped root bead. DC-ve current should be used. Pulsed current will give good control and a neat appearance when welding positionally.

As mentioned above, a large a ceramic shroud equipped with a gas lens is recommended to give the most effective gas shield and the arc length should be kept short; 3.5-4.5mm.  Argon or argon with small amounts of hydrogen, (1- 5%) are the appropriate shield gases with the Ar/H mixtures providing higher heat input. Above about 6mm thickness, TIG welding is generally replaced by the higher deposition rate MIG process, although mechanised/automated systems such as orbital TIG are very cost effective. A root purge of argon is recommended when welding a TIG root run and the next couple of fill passes.

MIG (GMAW) welding is carried out using either pure argon or argon-helium mixtures;  particularly useful on thicker sections. As with TIG, pulsed current will give better weld quality and appearance than dip transfer when welding out of the flat position. The filler wire is relatively soft and low friction liners are essential. Intermediate wire feeders may be required if the welding is taking place some distance from the

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wire drive unit. The filler wire pack should be opened at the last moment and should be adequately protected from contamination when installed in the wire feeder.

MMA (SMAW) welding electrodes are available, generally with a basic flux coating and designed to operate on DC+ve. Whilst these electrodes do not require baking before use, they may be dried at around 250OC if they have absorbed any moisture. Damp electrodes will result in weld metal porosity, as will a long arc. Weaving should be restricted to 3-4 times the electrode diameter. 

Submerged arc welding (SAW) becomes cost effective over a thickness of about 12.5mm if the component can be manipulated to enable welding to take place in the flat position. Weld preparation would be similar to that used for MIG welding. MIG wires of up to 2.4mm in diameter may be used so welding currents need to be correspondingly low, 300-350amps. The choice of welding flux should be discussed with the consumable supplier as an incorrect choice can result in slag detachability problems. 

Post weld heat treatment is not necessary but if dimensional stability is important the component may be stress relieved at 350-450OC.

Titanium and titanium alloys

Weldability of materials

Job Knowledge

Titanium and its alloys are chosen because of the following properties:

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high strength to weight ratio; corrosion resistance; mechanical properties at elevated temperatures.Titanium is a unique material, as strong as steel but half its weight with excellent corrosion resistance. Traditional applications are in the aerospace and chemical industries. More recently, especially as the cost of titanium has fallen significantly, the alloys are finding greater use in other industry sectors, such as offshore.

The various types of titanium alloys are identified and guidance given on welding processes and techniques employed in fabricating components without impairing their corrosion, oxidation and mechanical properties or introducing defects into the weld.

Material types

Alloy groupings

There are basically three types of alloys distinguished by their microstructure:

Titanium - Commercially pure (98 to 99.5% Ti) or strengthened by small additions of oxygen, nitrogen, carbon and iron. The alloys are readily fusion weldable.

Alpha alloys - These are largely single-phase alloys containing up to 7% aluminium and a small amount (< 0.3%) of oxygen, nitrogen and carbon. The alloys are fusion welded in the annealed condition.

Alpha-beta alloys - These have a characteristic two-phase microstructure formed by the addition of up to 6% aluminium and varying amounts of beta forming constituents - vanadium, chromium and molybdenum. The alloys are readily welded in the annealed condition.

Alloys which contain a large amount of the beta phase, stabilised by elements such as chromium, are not easily welded.

Commonly used alloys are listed in Table 1 with the appropriate ASTM grade, the internationally recognised designation. In industry, the most widely welded titanium alloys are the commercially pure grades and variants of the 6% Al and 4%V alloy.

Table 1: Commonly used titanium alloys and the recommended filler material

ASTM Grade Composition UTS (min) Mpa Filler Comments

1 Ti-0.15O 240 ERTi-1 Commercially pure

2 Ti-0.20O 340 ERTi-2 ,,

4 Ti-0.35O 550 ERTi-4 ,,

7 Ti-0.20O -0.2Pd 340 ERTi-7 ,,

9 Ti-3Al-2.5V 615 ERTi-9 Tube components

5 Ti-6Al-4V 900 ERTi-5 'Workhorse' alloy

23 Ti-6Al-4V ELI 900 ERTi-5ELI Low interstitials

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25 Ti-6Al-4V-0.06Pd 900 ERTi-25 Corrosion resistant grade

Filler alloys

Titanium and its alloys can be welded using a matching filler composition; compositions are given in The American Welding Society specification AWS A5.16-2004. Recommended filler wires for the commonly used titanium alloys are also given in Table 1.

When welding higher strength titanium alloys, fillers of a lower strength are sometimes used to achieve adequate weld metal ductility. For example, an unalloyed filler ERTi-2 can be used to weld Ti-6Al-4V and Ti-5Al-2.5Sn alloys in order to balance weldability, strength and formability requirements.

Weld imperfections

This material and its alloys are readily fusion welded providing suitable precautions are taken. TIG and plasma processes, with argon or argon-helium shielding gas, are used for welding thin section components, typically <10mm. Autogenous welding can be used for a section thickness of <3mm with TIG, or <6mm with plasma. Pulsed MIG welding using novel coated wires results in very low porosity and spatter.

The most likely imperfections in fusion welds are:

Weld metal porosity Embrittlement Contamination crackingNormally, there is no solidification cracking or hydrogen cracking.

Weld metal porosity

Weld metal porosity is the most frequent weld defect. Porosity arises when gas bubbles are trapped between dendrites during solidification. In titanium, hydrogen from moisture in the arc environment or contamination on the filler and parent metal surface, is the most likely cause of porosity.

It is essential that the joint and surrounding surface areas are cleaned by first degreasing either by steam, solvent, alkaline or vapour degreasing. Any surface oxide should then be removed by pickling (HF-HNO3 solution), light grinding or scratch brushing with a clean, stainless steel wire brush. On no account should an ordinary steel brush be used. After wiping with a lint-free cloth, care should be taken not to touch the surface before welding. When TIG welding thin section components, the joint area should be dry-machined to produce a smooth surface finish.

Embrittlement 

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Embrittlement can be caused by weld metal contamination by either gas absorption or by dissolving contaminants such as dust (iron particles) on the surface. At temperatures above 500°C, titanium has a very high affinity for oxygen, nitrogen and hydrogen. The weld pool, heat affected zone and cooling weld bead must be protected from oxidation by an inert gas shield (argon or helium). 

When oxidation occurs, the thin layer of surface oxide generates an interference colour. The colour can indicate whether the shielding was adequate or an unacceptable degree of contamination has occurred. A silver or straw colour shows satisfactory gas shielding was achieved but for certain service conditions, dark blue may be acceptable. Light blue, grey and white show a higher, usually unacceptable, level of oxygen contamination.

For small components, an efficient gas shield can be achieved by welding in a totally enclosed chamber, filled with the shielding gas. It is recommended that before welding, the arc is struck on a scrap piece of titanium, termed 'titanium-getter', to remove oxygen from the atmosphere; the oxygen level should be reduced to approximately 40ppm before striking the arc on the scrap titanium and <20ppm before welding the actual component.

In tube welding, a fully enclosed head is equally effective in shielding the weld area and is be preferable to orbital welding equipment in which the gas nozzle must be rotated around the tube.

When welding out in the open, the torch is fitted with a trailing shield to protect the hot weld bead whilst cooling. The size and shape of the shield is determined by the joint profile whilst its length will be influenced by welding current and travel speed. It is essential in 'open air' welding that the underside of the joint is protected from oxidation. For straight runs, a grooved bar is used with argon gas blown on to the joint. In tube and pipe welding, normal gas purging techniques are appropriate.

Contamination cracking

If iron particles are present on the component surface, they dissolve in the weld metal reducing corrosion resistance and, at a sufficiently high iron content, causing embrittlement. Iron particles are equally detrimental in the HAZ where local melting of the particles form pockets of titanium - iron eutectic. Microcracking may occur but it is more likely that the iron-rich pockets will become preferential sites for corrosion.

Particular attention should be paid to separating titanium from steel fabrications, preferably by designating a specially reserved clean area. Welders should guard against embedding steel particles into the surface of the material by:

Avoiding steel fabrication operations near titanium components.  Covering components to avoid airborne dust particles settling on the surface

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Not using tools, including wire brushes, previously used for steel Scratch brushing the joint area immediately before welding Not handling the cleaned component with dirty gloves.To avoid corrosion cracking, and minimise the risk of embrittlement through iron contamination, it is best practice to fabricate titanium in a specially reserved clean area.

Welding of titanium and its alloys - Part 1

Job Knowledge

Titanium is a reactive metal; it will burn in pure oxygen at 600°C and in nitrogen at around 800°C. Oxygen and nitrogen will also diffuse into titanium at temperatures above 400°C raising the tensile strength but embrittling the metal. In the form of a powder or metal shavings titanium also constitutes a fire hazard.

Despite this reactivity titanium is used extensively in chemical processing, offshore and aerospace applications. This is due to:

The tenacious protective oxide film that forms, giving the alloys very good corrosion resistance, particularly in chloride containing environments.

No loss of toughness at temperatures down to -196°C Good creep and oxidation resistance at temperatures up to almost 600°C. Similar strength to steel but at approximately half the weight.Because of the affinity of titanium and its alloys for oxygen, nitrogen and hydrogen and the subsequent embrittlement, fluxed welding processes are not recommended although they have been used, primarily in the former USSR. Arc welding is therefore restricted to the gas shielded processes (TIG, MIG and plasma-TIG) although power beams, the solid phase processes and resistance welding are also used.Titanium is allotropic; it has two different crystallographic forms depending on the temperature and chemical composition. Below 880°C it forms the hexagonal close packed alpha phase, above 880°C it exists as body centred cubic beta phase.

A range of elements may be used to improve the mechanical properties, some stabilise the alpha phase and others promote the formation of beta. Oxygen, carbon, nitrogen and aluminium promote the formation of the alpha phase; chromium, molybdenum, niobium, tin and vanadium promote the formation of beta. By suitable additions of these elements it is possible to produce four families of titanium alloys, divided on the basis of microstructure, into commercially pure titanium, alpha or near alpha alloys, alpha-beta alloys and beta alloys. ASTM designations, a simple numbering system, are a commonly used shorthand way of identifying the various alloys and both these and the alloy composition eg Ti-6Al-4V, will be used within this article.

Commercially pure, unalloyed ASTM 1 - 4 and 7 grades contain small amounts of contaminants such as oxygen, nitrogen and carbon, typically less than 0.2%, and have mechanical properties matching those of a good quality low carbon steel. The fewer contaminants, the lower is the tensile strength. The majority of these alloys are used for their corrosion resistance. Welding is straightforward and has little effect on the mechanical properties in the HAZ and they are generally welded in the annealed condition.

The alpha and near alpha alloys, typified by the Ti-5Al-2.5Sn alloy, have ultimate tensile strengths (UTSs) of 500-900MPa, 0.2% proof (PS) of 600-800MPa and elongations (El%) of around 18%. As with the commercially pure alloys the mechanical properties of this group are insensitive to heat treatment. Weldability is good, the alloys being welded in the annealed condition.

The alpha-beta alloys are sensitive to heat treatment, solution treatment and ageing, increasing the strength by 50% compared with the annealed condition. The very high strength alpha-beta alloys such as Ti-5Al-2Sn-2Zr-4Mo-4Cr may have a UTS of 1200MPa, PS of 1150MPa and an El% of 10. Weldability of the

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alloys within this group, however, is dependent on the amount of beta present, the most strongly beta stabilised alloys being embrittled during welding and, although it is possible to restore some of the ductility by a post-weld heat treatment, this is often impractical. These very high strength, high beta content alloys are therefore rarely welded. Contrast this with possibly the most frequently used alpha-beta alloy, Ti-6Al-4V (ASTM Grade 5) with a UTS of 950MPa, a PS of 850MPa and El% of 15. This alloy has good formability, is readily workable, has good castability, excellent weldability and could be regarded as the alloy against which to benchmark all others.

The fully beta alloys, eg Ti-13V-11Cr-3Al, have similar strengths but with slightly improved ductility, typically around 15% elongation. The beta phase is termed metastable - cold work or heating to elevated temperatures may cause partial transformation to alpha. The alloys have high hardenability, very good forgeability and are very ductile. Weldability is good, taking place with the alloy in the annealed or solution treated condition although to obtain the full strength it is generally necessary to weld in the annealed condition, cold work, solution treat and then carry out an ageing treatment.

Filler metals, all solid wires and matching the composition of the commoner of the alloys, are available, the relevant specifications being AWS A5.16/A5.16M:2007 Specification for titanium and titanium-alloy welding electrodes and rods and BS EN ISO 24034.2010 Welding consumables, solid wires and rods for fusion welding of titanium and titanium alloys. Although readily available, the range of consumables is somewhat restricted with only fourteen or fifteen compositions being produced in accordance with these specifications.

Weldability, as mentioned above, is in general very good. The exception is the high beta alpha-beta alloys. The fundamental problem in welding titanium alloys is the elimination of atmospheric contamination. Contamination of the weld metal and the adjacent HAZs will increase tensile strength and hardness but may reduce ductility to an unacceptably low value such that cracks may occur even in conditions of only moderate restraint. The most likely contaminants are oxygen and nitrogen, picked up due to air entrained in the gas shield or from impure shield gas, and hydrogen from moisture or surface contamination.

TIG welds in commercially pure titanium sheet made with successively greater air contamination of the shielding

The maximum tolerable limits in weld metal have been estimated as 0.3% oxygen, 0.15% nitrogen and 150ppm hydrogen so scrupulous cleanliness is essential for both parent metals and filler wires. Degreasing

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the weld preparation followed by stainless steel wire brushing and a further degrease is generally sufficient. Heavily oxidised components may need to be pickled in a nitric/hydrofluoric acid mixture to remove the oxide layer. Degreasing of the filler wire for TIG welding should be done as a matter of course and the cleaned wire handled with clean cotton gloves; grease and perspiration from the fingers can cause local contamination and/or porosity. MIG wire should be ordered in a degreased condition, stored in clean dry conditions and not left unprotected on the shop floor.

During welding those parts of the weldment exposed to temperatures above 520°C will absorb oxygen and nitrogen and must therefore be protected until they have cooled below this critical temperature. Fortunately heat conduction in titanium is low so the area affected is limited in size and chill blocks can be used to reduce this heated zone even further. The molten weld pool is protected by the normal gas shroud but the cooling weld and its HAZ will need additional protection by the use of a trailing shield with its own protective gas supply following along behind the welding torch. The back face of the weld also needs similar protection by the provision of an efficient gas purge.

Surface discolouration will give a good indication of the degree of atmospheric contamination as shown in the the colour chart. Under perfect shielding conditions the weld will be bright and silvery in appearance. Discolouration at the outer edges of the HAZ is not generally significant and may be ignored. As contamination increases the colour changes from silver to a light straw colour, then dark straw, dark blue, light blue, grey and finally a powdery white.

The light and dark straw colours indicate light contamination that is normally acceptable. Dark blue indicates heavier contamination that may be acceptable depending on the service conditions. Light blue, grey and white indicate such a high level of contamination that they are regarded as unacceptable. In multi-pass welds the contamination will obviously affect any subsequent weld runs so that surface appearance alone is not a reliable guide to whether or not unacceptable contamination has occurred. A simple bend test is a reliable but destructive method of checking if the weld is unacceptably embrittled but note that the bend radius varies depending on the particular alloy. For example, a 3t bend radius is used for testing a Grade 2 weld but a 10t bend radius is used when testing Ti-6Al-4V. Portable hardness checks may also be carried out on production items; this requires knowledge of the hardness expected in the specific alloy weld metal.

Part 2 of this article will consider some of the other welding problems and provide guidance on TIG and MIG welding of titanium.

Welding of titanium and its alloys - Part 2

Job Knowledge

Part 1

Courtesy Huntingdon Fusion Techniques

Titanium and its alloys are remarkably resistant to the cracking problems experienced by many of the other alloy systems. Solidification and liquation cracking are virtually unknown and what could perhaps be called cold cracking, occurs generally only because of embrittlement arising from contamination, as discussed

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in Part 1.

Porosity is the commonest problem, particularly when close square butt joints are used. It is generally attributed to hydrogen and cleanliness is therefore crucial in eliminating porosity. The porosity may be of one or a mixture of two types: firstly micro-porosity formed within the arms of the dendrites during solidification and secondly, larger pores that often align themselves along the weld centre line.

As discussed in Part 1, cleanliness is the key to defect free welds and this means that not only must the component be thoroughly degreased but so should the filler wires; weld preparation edges must be deburred and the highest purity shielding gas must be used. Ideally the gas should have a dew point of less than -50°C (39ppm H2O) and to maintain this low level the gas supply system should be free of leaks. Regular and frequent maintenance of the system is therefore essential, checking the joints for leaks and for damaged hoses. Ideally the gas supply should befrom a bulk gas tank, not cylinders, and delivered to the work stations via welded or brazed steel or copper tubing. Plastic hoses should be kept as short as possible; most plastics used are porous and will allow moisture to permeate through the hose wall; neoprene and PVC are the worst, Teflon one of the least porous. It is worth remembering that moisture can collect in the hose over a period of time so a porosity problem, say after a weekend shut down, may be an indication that this is occurring.

TIG filler wires should be cleaned with a lint free cloth and an efficient degreasant immediately before use. Following cleaning, the wire should not be handled with bare hands but whilst wearing clean, grease-free gloves. MIG wire presents more of a problem but devices to clean the wire as it passes through the wire feeder are available. For the best results wire that has been shaved to remove any embedded contaminants can be obtained.

A further potential source of contamination that is frequently overlooked is the use of air powered tools for wire brushing or dressing weld preparations and welds. Most compressed air contains moisture and oil so that, even when oil and moisture traps are fitted, it is possible to leave a thin film of moisture and/or oil on the surface to be welded. It is recommended that electrically powered tools are used at all times once the item has been degreased prior to welding.

Although regarded as a very minor problem, ductility dip cracking (where alloys experience a severe loss of ductility at a temperature below the solidification temperature) has been noted in some of the titanium alloys; the alpha-beta alloys containing niobium being the most susceptible with Ti-6Al-2Nb-1Ta-0.8Mo the most sensitive. The temperature range in which this loss of ductility occurs is between750°C and 850°C.

The cracking is intergranular and is thought to be partly the result of volume changes during the beta to alpha phase change coupled with the reduction in ductility.

A significant amount of welding of titanium alloys is carried out without the use of filler metals. When filler wire is used, generally a composition matching that of the parent metal is selected. There are, however, some exceptions. The welding of high strength but low ductility commercial purity titanium is generally performed with a low strength filler metal in order to achieve the desired weld quality. Similarly, unalloyed filler metal is sometimes used to weld alloys such as Ti-6Al-4V, thereby improving weld metal ductility by lowering the amount of beta phase that is formed. Extra low interstitial (ELI) filler metals are also available and may be used to improve weld metal ductility and toughness.

Most of the titanium alloys can be successfully fusion welded using the gas shielded welding processes and power beams; all can be welded using solid phase processes, friction and resistance welding. Welding parameters and weld preparations are similar to those that would be used to weld a carbon steel. From the

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welder's point of view, titanium iseasier to weld than steel, having good fluidity and high surface tension, easing the task of depositing sound full penetration root beads.

TIG welding is probably the most commonly used process in both manual and mechanised applications. The current is DC-ve, generally with high purity argon as the shielding gas, although helium or Ar/He mixtures may be used to improve penetration. Torch nozzles should be fitted with gas lenses to improve gas shielding and the ceramic shroud should be as large a diameter as possible. A 1.5mm diameter tungsten, for example, should be used with a 16mm diameter ceramic. Arc lengths need to be as short as possible to reduce the risk of contamination; 1 to 1.5 times the electrode diameter is regarded as a good rule of thumb. Arc initiation should be achieved by the use of HF current or Lift Arc to prevent tungsten contamination. The equipment must also be capable of continuing the shield gas flow after the arc is extinguished so that the weld can cool within the protective gas shield. It is also advisable to keep the tip of the filler wire within the gas shield until such times as it has cooled to a sufficiently low temperature.

A supplementary trailing gas shield will also need to be attached to the torch to provide protection to the cooling weld metal as the welder moves along the joint line. This makes manipulation of the welding torch more difficult. Most welders manufacture their own supplementary shields, shaped to closely fit the component; several shields would therefore be required to weld a range of pipe diameters. A backing gas is also necessary and back purging should be maintained for at least the first three or four passes in a weld. Backing gas purity should be better than 20ppm maximum oxygen.

MIG welding using argon or argon/helium mixtures may be used but this process will not provide the same high quality weld metal as the TIG process and it can be difficult to achieve the stringent quality levels required by aerospace applications. Dip transfer can lead to lack of fusion defects and spray transfer requires both leading and trailing supplementary gas shields, the leading gas shield to prevent oxidation of any spatter that may be remelted into the weld pool. The improvements in pulsed MIG power sources by the use of inverter technology and micro-processor control have obviated some of these problems and substantially narrowed the gap between MIG and TIG. MIG is, however, still difficult for the manual welder because of the difficulty of manipulating the MIG torch with a supplementary gas shroud. Because of these difficulties MIG welding is often mechanised or automated.

Plasma-TIG may be used for welding titanium, being capable of keyholing a weld up to 12.5mm thick. The same requirements for gas purity and weld pool protection required for TIG are also needed for plasma-TIG. Plasma-TIG is rarely used in a manual application and never in the keyhole mode.

Atmospheric contamination is best avoided by the use of a welding chamber or glove box that can be filled with argon. Purpose built glove boxes can be purchased but it is a simple matter to fabricate a chamber of an appropriate size using slotted angle eg DexionTM angle, to form the frame and covering this with a clear plastic or acetate sheet. The size of the component that can be welded within a glove box is necessarily restricted.

Electron beam, laser, friction, resistance spot and seam and flash welding are all used to weld titanium and its alloys. Electron beam welding, being carried out in a vacuum, needs no protective gas shield. Conventional friction welding may also be carried out without a protective shield although a gas shield should be used when friction stir welding. Similarly, no gas shield is required when resistance welding, although for the most critical applications a gas shield is recommended. Laser and flash welding both require gas shielding for the best results and least atmospheric contamination.

Cast irons

Job Knowledge

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Weldability of materials

Cast irons are iron based alloys containing more than 2% carbon, 1 to 3% silicon and up to 1% manganese. As cast irons are relatively inexpensive, very easily cast into complex shapes and readily machined, they are an important engineering and structural group of materials. Unfortunately not all grades are weldable and special precautions are normally required even with the so-called weldable grades.

Material types

Cast irons can be conveniently grouped according to their structure which influences their mechanical properties and weldability; the main groups of general engineering cast irons are shown in Fig. 1.

Fig.1. Main groups of engineering cast irons

Grey cast irons

Grey cast irons contain 2.0 - 4.5% carbon and 1 - 3% silicon. Their structure consists of branched and interconnected graphite flakes in a matrix which is pearlite, ferrite or a mixture of the two (Fig.2a). The graphite flakes form planes of weakness and so strength and toughness are inferior to those of structural steels. 

Nodular cast irons

The mechanical properties of grey irons can be greatly improved if the graphite shape is modified to eliminate planes of weakness. Such modification is possible if molten iron, having a composition in the range 3.2 - 4.5% carbon and 1.8 - 2.8% silicon, is treated with magnesium or cerium additions before casting. This produces castings with graphite in spheroidal form instead of flakes, known as nodular, spheroidal graphite (SG) or ductile irons (Fig.2b). Nodular irons are available with pearlite, ferrite or pearlite-ferrite matrices which offer a combination of greater ductility and higher tensile strength than grey cast irons.

Fig.2. Microstructures of a) grey cast iron and

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b) spheroidal graphite cast iron

White cast irons

By reducing the carbon and silicon content and cooling rapidly, much of the carbon is retained in the form of iron carbide without graphite flakes. However, iron carbide, or cementite, is extremely hard and brittle and these castings are used where high hardness and wear resistance is needed.

Malleable irons

These are produced by heat treatment of closely controlled compositions of white irons which are decomposed to give carbon aggregates dispersed in a ferrite or pearlitic matrix. As the compact shape of the carbon does not reduce the matrix ductility to the same extent as graphite flakes, a useful level of ductility is obtained. Malleable iron may be divided into classes. Whiteheart, Blackheart and Pearlitic irons.

Whiteheart malleable irons

Whiteheart malleable castings are produced from high carbon white cast irons annealed in a decarburising medium. Carbon is removed at the casting surface, the loss being only compensated by the diffusion of carbon from the interior. Whiteheart castings are inhomogenous with a decarburised surface skin and a higher carbon core.

Blackheart malleable irons

Blackheart malleable irons are produced by annealing low carbon (2.2 - 2.9%) white iron castings without decarburisation. The resulting structure, of carbon in a ferrite matrix, is homogenous with better mechanical properties than those of whiteheart irons.

Pearlitic malleable irons

These have a pearlitic rather than ferritic matrix which gives them higher strength but lower ductility than ferritic, blackheart irons.

Weldability

This depends on microstructure and mechanical properties. For example, grey cast iron is inherently brittle and often cannot withstand stresses set up by a cooling weld. As the lack of ductility is caused by the coarse graphite flakes, the graphite clusters in malleable irons, and the nodular graphite in SG irons, give significantly higher ductility which improves the weldability.

The weldability may be lessened by the formation of hard and brittle microstructures in the heat affected zone (HAZ), consisting of iron carbides and martensite. As nodular and malleable irons are less likely to form martensite, they are more readily weldable, particularly if the ferrite content is high.

White cast iron which is very hard and contains iron carbides, is normally considered to be unweldable.

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Welding process

Braze welding is frequently employed to avoid cracking. Braze welding is often called 'Bronze welding' in the UK. Bronze welding is a varient of braze welding employing copper-base fillers, it is regulated by BS 1724:1990. (This standard has been withdrawn, but no direct replacement has been identified.) As oxides and other impurities are not removed by melting, and mechanical cleaning will tend to smear the graphite across the surface, surfaces must be thoroughly cleaned, for example, by means of a salt bath.

In fusion welding, the oxy-acetylene, MMA, MIG/FCA welding processes can all be used. In general, low heat inputs conditions, extensive preheating and slow cooling are normally a pre-requisite to avoid HAZ cracking.

Oxy-acetylene because of the relatively low temperature heat source, oxy-acetylene welding will require a higher preheat than MMA. Penetration and dilution is low but the wide HAZ and slow cooling will produce a soft microstructure. Powder welding in which filler powder is fed from a small hopper mounted on the oxy-acetylene torch, is a very low heat input process and often used for buttering the surfaces before welding.

MMA widely used in the fabrication and repair of cast iron because the intense, high temperature arc enables higher welding speeds and lower preheat levels. The disadvantage of MMA is the greater weld pool penetration and parent metal dilution but using electrode negative polarity will help to reduce the HAZ.

MIG and FCA MIG (dip transfer) and especially the FCA processes can be used to achieve high deposition rates whilst limiting the amount of weld penetration.

Filler alloys

In oxy-acetylene welding, the consumable normally has slightly higher carbon and silicon content to give a weld with matching mechanical properties. The most common MMA filler rods are nickel, nickel - iron and nickel - copper alloys which can accommodate the high carbon dilution from the parent metal and produces a ductile machinable weld deposit.

In MIG welding, the electrode wires are usually nickel or Monel but copper alloys may be used. Flux cored wires, nickel-iron and nickel-iron-manganese wires, are also available for welding cast irons. Powders are based on nickel with additions of iron, chromium and cobalt to give a range of hardnesses.

Weld imperfections

The potential problem of high carbon weld metal deposits is avoided by using a nickel or nickel alloy consumable which produces finely divided graphite, lower porosity and a readily machinable deposit. However, nickel deposits which are high in sulphur and phosphorus from parent metal dilution, may result in solidification cracking.

The formation of hard and brittle HAZ structures make cast irons particularly prone to HAZ cracking during post-weld cooling. HAZ cracking risk is reduced by preheating and slow post-weld cooling. As preheating will slow the cooling rate both in weld deposit and HAZ, martensitic formation is suppressed and the HAZ hardness is somewhat reduced. Preheating can also dissipate shrinkage stresses and reduce distortion, lessening the likelihood of weld cracking and HAZ.

Table 1: Typical preheat levels for welding cast irons

Cast iron type Preheat temperature degrees C

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MMA MIG Gas (fusion) Gas (powder)

Ferritic flake 300 300 600 300

Ferritic nodular RT-150 RT-150 600 200

Ferritic whiteheart malleable RT* RT* 600 200

Pearlitic flake 300-330 300-330 600 350

Pearlitic nodular 200-330 200-330 600 300

Pearlitic malleable 300-330 300-330 600 300

RT - room temperature* 200 degrees C if high C core involved.

As cracking may also result from unequal expansion, especially likely during preheating of complex castings or when preheating is localised on large components, preheat should always be applied gradually. Also, the casting should always be allowed to cool slowly to avoid thermal shock.

An alternative technique is 'quench' welding for large castings which would be difficult to preheat. The weld is made by depositing a series of small stringer weld beads at a low heat input to minimise the HAZ. These weld beads are hammer peened whilst hot to relieve shrinkage stresses and the weld area is quenched with an air blast or damp cloth to limit stress build up.

Repair of castings

Because of the possibility of casting defects and their inherent brittle nature, repairs to cast iron components are frequently required. For small repairs, MMA, oxy-acetylene, braze and powder welding processes can all be used. For larger areas, MMA or powder technique can be used for buttering the edges of the joint followed by MMA or MIG/FCA welding to fill the groove. This is shown schematically in figure 3.

Fig.3. Repair of crack in cast iron from one side

  a) bridging the crack by weld bead from buttered layers 

b) sequence of welding

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Remove defective area preferably by grinding or tungsten carbide burr. If air arc or MMA gouging is used, the component must be preheated locally to typically 300 degrees C.

After gouging, the prepared area should be lightly ground to remove any hardened material. Preheat the casting to the temperature given in Table 1.  Butter the surface of the groove with MMA using a small diameter (2.4 or 3mm) electrode; use a

nickel or Monel rod to produce a soft, ductile 'buttered' layer; alternatively use oxy-acetylene with a powder consumable.

Remove slag and peen each weld bead whilst still hot. Fill the groove using nickel (3 or 4mm diameter) or nickel-iron electrodes for greater strength.Finally, to avoid cracking through residual stresses, the weld area should be covered to ensure the casting will cool slowly to room temperature.

Welding of HSLA Steels

Job Knowledge

The development and use of high strength low alloy (HSLA) steels has been driven by the need to reduce costs, the higher strength compared with a conventional carbon-manganese steel enabling thinner and lighter structures to be erected. The majority of these steels are to be found in structural applications; offshore structures, yellow goods, buildings, shipbuilding etc. Tensile strengths of up to 690MPa are achievable whilst still maintaining good weldability and high notch toughness, often better than 50J at -60°C.

There are two methods by which both high tensile strength and toughness is achieved - by micro-alloying, adding small amounts of strong carbide and nitride formers and by very careful control of the rolling temperature - controlled rolling or thermo-mechanically controlled processing (TMCP steels).

The highest strengths are achieved by a combination of the two methods. The aim of both methods is to produce as small a grain size as possible, fine grain giving the best notch toughness and each halving of the grain diameter producing a 50% increase in tensile strength.

Improved weldability is an additional objective and this is achieved by reducing the hardenability of the steel, the carbon content of some steels being lower than 0.05%C, and reducing undesirable elements such as sulphur and phosphorous to as low a level as possible.

To compensate for the loss of carbon and to increase tensile strength small additions of alloying elements such as niobium (<0.10%), titanium (<0.030%) and vanadium (<0.15%) are made, perhaps also with small amounts of molybdenum, chromium, copper and nitrogen. These elements are strong carbide and nitride

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formers, producing a fine dispersion of stable precipitates that inhibit grain growth during hot rolling and assist in nucleating fine grained ferrite during cooling.

These elements also provide some increase in strength by precipitation hardening. Controlled rolling by the TMCP hot rolling method may also be used to provide additional grain refinement and hence an increase in tensile strength and toughness. TMCP is carried out at a temperature about or just below the recrystallisation temperature of the steel i.e.below about 900°C, resulting in elongated crystals of austenite. Accelerated cooling from the rolling temperature then causes very fine grained ferrite to form on the austenite grain boundaries.

Despite the improved weldability of these steels there are some fabrication problems. Firstly, hydrogen induced cold cracking.

The low carbon content - and hence low carbon equivalent, sometimes less than 0.30CEv - means that these steels have a low sensitivity to hydrogen cold cracking (see Job Knowledge 45 but note that the standard IIW carbon equivalent formula is not valid for all of these steels and cannot always be relied upon when calculating preheat temperatures).

The HSLA steels can therefore be welded with lower preheats than would be permitted for conventional carbon-manganese steels, despite their higher strength. The highest risk of cold cracking in these types of steels is therefore in the weld metal, rather than the HAZ. There are several reasons for this; a) The high strength of the parent metal means higher residual stresses during welding, b) To match the tensile strength and toughness of the parent steel, the filler metals need to be more highly alloyed and therefore will have a higher CEv, perhaps as high as 0.6CEv (IIW) if matching the tensile strength of a 700MPa yield steel with an E11018-G electrode. c) The weld metal transforms from austenite to ferrite at a lower temperature than the parent steel (it is generally the other way round in a conventional carbon-manganese steel) meaning that any hydrogen in the HAZ is rejected into the still austenitic weld metal which has a high solubility for hydrogen. A preheat based on the weld metal composition is therefore advisable and low hydrogen techniques must be used. The exceptions to this rule are those HSLA pipeline steels specifically designed to be welded with cellulosic electrodes. Advice regarding the preheat temperature for specific steels should be sought from the steel manufacturer.

Secondly, even though steels generally have very low levels of sulphur, the steels containing less than 0.05%C may suffer from solidification cracking in the root pass of butt joints, particularly if the root bead is deposited at a high welding speed. The reason for this is that high dilution of the filler metal produces a weld metal low in carbon. This low carbon content in its turn leads to excessive grain growth of the austenite during welding and these large grains increase the risk of centre line solidification cracking in the root bead. This problem appears to be most prevalent in pipe butt joints welded using cellulosic electrodes, probably due to it being possible to use a fast, vertical-down welding technique.

Thirdly, toughness and strength in the HAZ can be an issue. The steel manufacturer takes great care to control rolling temperatures and cooling rates to provide the desired properties. The component is then welded, producing a heat affected zone that has experienced an uncontrolled cycle of heat treatment. The microstructure in the HAZ will vary with respect to the composition of the steel and the welding process heat input. A high heat input will promote grain growth and this will have an adverse effect on both strength and toughness. As a rule of thumb, heat input should be restricted to around 2.5kJ/mm maximum and the interpass temperature maintained at 250°C maximum, although some of the steels containing titanium and boron can tolerate heat inputs as high as 4.5kJ/mm without undue loss of strength. For a definitive statement on heat input control the advice of the steel manufacturer should be sought.

These steels must under no circumstances be normalised or tempered although post weld heat treatment (PWHT) is often a requirement when the component thickness is greater than some 35 to 40mm. Care needs

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to be taken if PWHT is applied that the soak temperature does not exceed 600°C; a temperature range of 550°C to 600°C is often specified. The reason for this is that many of the TMCP steels are accelerated cooled to a temperature of around 620°C; heat treating at or close to this temperature will result in a substantial reduction in tensile strength due to over-tempering. The same restriction applies to any hot working activity - plate must not be hot rolled and the temperature of local heating for correction of distortion must not be allowed to exceed 600°C.

Welding of ferritic creep-resistant steels

Job Knowledge

Creep is a long term failure mechanism that, in most metals, occurs at elevated temperatures (see Job Knowledge No. 81   ). Creep strength in the ferritic steels is achieved by alloying with elements that will provide enhanced strength at high temperatures. Chromium (Cr) and molybdenum (Mo) are the two principal alloying elements but vanadium (V) and niobium (Nb) may also be added.

Table 1 gives the nominal composition of the commoner creep resistant steels. In addition to the use of these steels in creep service they also have resistance to hydrogen attack and corrosion by sulphur bearing hydrocarbons. They are therefore found in power generation and the oil and gas industries.

Table 1 Nominal composition and mechanical properties of the creep resistant steels

Steel grade Composition - nominal % Mechanical properties - typical

  C max Cr Mo V Nb UTSN/mm2

0.2% ProofN/mm2

Elongn %

Charpy-VJ/°C min

C1/2Mo 0.3   0.5     520 320 35 27@20½/½/¼CrMoV 0.14 0.5 0.6 0.25   585 330 25  1¼Cr½Mo 0.13 1.2 0.5     530 350 35 27@202¼Cr1Mo 0.13 2.25 1.0     555 350 35 27@205Cr½Mo 0.13 5.0 0.5     690 475 28 27@239Cr1Mo 0.13 9.0 1.0     675 475 30  9CrMoVNb (9Crmod or P91)

0.13 8.75 1.0 0.23 0.08 650 480 30 40@20

The creep resistant steels all contain strong carbide and/or nitride forming elements. These are intended to provide a fine dispersion of precipitates that both increase the tensile strength and impede the formation of the voids illustrated in Fig 1 and Fig 2 of Job Knowledge No. 81. Chromium is also added to reduce the scaling or oxidation of the steel at high temperatures. Each steel grade has a creep limit (a stress and temperature above which it should not be used) and a similar limit on oxidation resistance. The allowable temperature increases with the alloy content, enabling the more highly alloyed steels to be used up 650°C.

The ½/½/¼CrMoVsteel is a special case. It was developed for the power generation industry in the UK and is unlikely to be encountered elsewhere but some notes have been included as it may be found in older plant scheduled for repair.

As the alloy content increases then so does the hardenability (the ability to form martensite) of the steel. C½Mo, CMV and 1¼Cr½Mo steels form ferritic/bainitic structures, the other more highly alloyed steels forming martensite, even at relatively slow cooling rates. This should give some hint as to one of the problems encountered when welding this family of steels; that of hydrogen induced cold cracking (see Job Knowledge No. 45), since martensite is generally hard, brittle and sensitive to the presence of hydrogen.

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Low hydrogen welding processes are therefore essential. This includes ensuring that any shield gases are of high purity and are dry; ideally with a dew point less than 50°C.

Preheat is essential for most of the alloys (the IIW carbon equivalent method is not valid for these grades of steel) and few welding specifications give much guidance regarding recommended preheat temperatures. However, ASME B31.3 and EN 1011 Part 2 both contain recommendations. Table 2 is adapted from the EN specification for processes with hydrogen limited to between five and 10mls of hydrogen in 100gms of weld metal (Scale C). It may be permissible to use lower preheats if the hydrogen content is reduced to less than 5mls/100gm; for instance when depositing a TIG root pass. This could be confirmed during welding procedure development.

Table 2 Recommended preheat and interpass temperatures

Steel Grade Thickness(mm)

Min. Preheat(°C)

Max. Interpass(°C)

C1/2Mo≤15 >15≤30 >30

20 75 100

250

½/½/¼CrMoV All 150 300

1¼Cr½Mo ≤15 >15

100 150 300

2¼Cr1Mo ≤15 >15

150 200 350

5Cr½Mo All 200 3509Cr1Mo All 200 3509Cr1MoVNb All 200 350

11/4Cr1/2Mo power station boiler header

An additional problem that may be encountered with the creep resistant steels is that of reheat cracking (see Job Knowledge No. 48   ). This is a cracking mechanism that takes place, as the name suggests, during reheating of the welded joint, either when the weld is post weld heat treated (PWHT) or is put into high temperature service without PWHT.

The most sensitive grades are those containing vanadium; the ½/½/¼CrMoV steel being one of the most sensitive. It is so sensitive that it may be necessary to maintain the preheat and hot grind and blend the weld

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toes of a thick, highly restrained weld to reduce stress concentrations before immediately performing the PWHT operation.

Solutions to this problem are control of residual elements to low levels, low heat input to minimise grain growth in the HAZ and devising a welding procedure that results in the maximum amount of grain refinement in the HAZ. Rapid heating through the temperature range 350 - 600°C at which the steel is most sensitive can also help. This approach must be treated with some caution as too rapid a temperature rise can cause unacceptable stresses and distortion and may violate code requirements.

Most of the creep resistant steels require PWHT; mandatory in all of the application codes. This is to ensure that the hard microstructures formed during welding are softened and toughness improved. It is also necessary to heat treat the weld and HAZs to ensure that the precipitates, required to give best creep performance, are of the correct size and distribution. PWHT temperatures and soak times must therefore be closely controlled to develop the required mechanical properties. Typical temperatures and times are given in Table 3. These figures are typical only and it is important that the item is heat treated precisely in accordance with the relevant application code; ASME VIII, BS PD5500, EN 13445 etc. 

Table 3 Typical PWHT temperatures and times

Steel Grade TemperatureRange (°C)

Soak Time(hours)

C1/2Mo 630 - 670 1 per 25mm½/½/¼CrMoV 650 - 680 1 per 25mm1¼Cr½Mo 650 -700 1 per 25mm2¼Cr1Mo 680 - 720 2 min5Cr½Mo 710 - 750 2 min9Cr1Mo 730 - 760 2 min9Cr1MoVNb 730 - 760 2 min

The PWHT temperature of the 1¼Cr½Mo and the 2¼Cr1Mo steels are sometimes changed from the ranges given inTable 3 in order to develop specific properties; see for example Table 4.4.1 in BS PD5500.

The 9CrMoVNb steel is particularly sensitive to PWHT times and temperatures and great care must be exercised when PWHT'ing this particular grade of steel.

Any alloy containing more than 2% chromium will need to be bore purged with an inert gas such as argon when depositing a TIG root pass. Exposure of the molten weld pool to the atmosphere results in some of the chromium boiling off giving rise to a porous or 'coked' bead on the reverse side of the weld. This adversely affects both mechanical properties and corrosion resistance.

Welding consumables matching the parent metal composition are readily available for all of these steels for most of the welding processes. An exception to this is the ½/½/¼CrMoV steel which is conventionally welded with a 2¼Cr1Mo filler. Dissimilar metal joints made between components from within this group of ferritic steels or with the carbon manganese steels are usually welded using a filler metal that matches the less highly alloyed steel. PWHT temperature for the dissimilar metal joints can be a problem and tends to be a compromise between overtempering the lower alloyed steel and undertempering the more highly alloyed metal.

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Welding of ferritic cryogenic steels

Job Knowledge

Ferritic cryogenic steels are nickel containing low alloy steels designed to operate safely at temperatures substantially below 0°C and are characterised by good tensile properties and high impact strength at low temperatures.

The nickel content ranges from around 1.5 to 9%, although there are some fine grained carbon-manganese steels that may be operated at temperatures as low as -50°C. These grades of steel are generally found in the oil and gas and petrochemical industries where they are used for the handling and storage of liquefied petroleum gases (LPG) at temperatures down to approximately -100°C and, in the case of the 9% nickel steel, down to -196°C. They are also found in the gas processing industry for the production and handling of gases such as carbon dioxide and oxygen as shown in Table 1.

Table 1. Approximate minimum service temperatures and applications of the cryogenic steels

Steel Type Specification(Plate)

Minimum service

temperature °C

Typical storage/processing application

Fine grained Al killed C/Mn steel

EN10028-3P460NL2 -50 Ammonia, propane (LPG)

1.5% Ni steel EN10028-4 15NiMn6 -60 Ammonia, propane, carbon disulphide

2.5% Ni steel ASTM A203 GrB -60 Ammonia, propane, carbon disulphide

3.5% Ni steel ASTM A203Gr EEN10028-4 12Ni14 -101 Carbon dioxide, acetylene, ethane

5% Ni steel EN10028-4 X12Ni5 -130 Ethylene (LEG)

9% Ni steelASTM A353/A553Tp1EN10028-4 X8Ni9

-196 Methane (LNG), oxygen, argon

Austenitic stainless steel ASTM 304LEN10088-1 1.4305 -273 Nitrogen, hydrogen, helium

The choice of which steel to use for any particular application depends not only on the temperature but also on such aspects as section thickness required by design and the possibility of stress corrosion.

The applications of these steels require that the mechanical properties, in particular the toughness, of welds and their associated heat affected zones match or are very close to those of the parent metals. The

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fabrication of the cryogenic steels into pipework and vessels therefore requires careful selection of welding consumables and close control of welding parameters.

Manual metal arc (MMA) electrodes matching the composition and Charpy-V impact strength of the fine grained carbon manganese steels at -50°C can be obtained, for example, AWS A5.5 E7018-1 electrodes, although the addition of a small amount of nickel, up to 1%, will give added confidence in achieving the required toughness. Matching C/Mn composition metal active gas (MAG), flux cored (FCAW) and submerged arc (SA) consumables will not give adequate toughness at -50°C and require nickel to provide the required as-welded toughness.

This is generally limited to a maximum of 1%Ni to comply with the NACE International ISO15156-2/MR0175 requirement for use in sour service. For even greater confidence that acceptable Charpy-V values can be achieved and to provide an improved tolerance to procedural variations then 2.5% nickel containing consumables may be used.

The 1.5%Ni and 2.5%Ni steels may be welded with 2.5% Ni consumables and these will provide adequate toughness down to -60°C in both the as-welded and post weld heat treated (PWHT) condition. A word of caution, however; the tensile strength of PWHT'd TIG and MAG weld metal may fall below the minimum specified for the parent metal. MAG weld metal deposited using a shield gas with a high proportion (>20%) of CO2 appears to be particularly sensitive.

Consumables are available for the MMA and SAW welding processes but not for the TIG, MAG or FCAW processes.

For depositing TIG root passes in the 3.5 Ni alloys, a 2.5% Ni filler metal is normally used. Although the 3.5% Ni consumables are capable of providing adequate toughness at -101°C they are very sensitive to variations in welding parameters, heat input and welding position. This sensitivity results in a wide variability of impact test results so for the more demanding applications, alternative nickel based filler metals such as AWS ENiCrFe-2 or EniCrFe-3 are often used enabling all of the conventional arc welding processes to be used.

The 5% Ni and 9% Ni alloys are conventionally welded using a nickel based filler metal. 6.5% Ni MMA electrodes are available but these are not capable of consistently providing adequate toughness much below -110°C. Consumables for welding the 9% Ni alloy have been developed; these typically contain 12% to 14% nickel. However, the cost of production is such that they do not compete with the nickel based alternatives.

A problem with the nickel based consumables that were initially used to weld these steels is that their tensile strength is substantially less than that of the parent metal. Higher strength fillers of the AWS EniCrMo-3 (alloy 625) type are now readily available and these enable all the arc welding processes to be used. They also match parent metals with respect to toughness and ultimate tensile strength although the 0.2% proof strength of TIG, MIG and SAW weld metals may fall below that specified for the 9%Ni steel.

As with any steel where good toughness is required, heat input must be controlled. It is recommended that interpass temperatures are limited to a maximum of 250°C and ideally less than 150°C for the 9%Ni alloy. Heat input from welding should be limited to approximately 3.5kJ/mm for SAW and 2.5kJ/mm for MMA.

Preheat may be required for the carbon-manganese and up to 3.5% Ni alloys, depending upon section thickness, joint type and restraint to reduce the risk of hydrogen cold cracking. ASME B31.3, for example, recommends minimum preheat temperatures of 79°C for carbon steels greater than 25mm thick, 93°C for all thicknesses of the 1.5%, 2.5% and 3.5% nickel steels but only 10°C for the 5% and 9% Ni alloys. The reason for this low preheat temperature is that these high nickel content alloys contain a large amount of

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austenite that can tolerate large amounts of hydrogen. This austenite therefore substantially reduces the risk of cold cracking; in addition, they are conventionally welded with nickel based alloys that reduce the risk even further.

Post weld heat treatment is not generally required for the 9%Ni steels; indeed, EN 13445-4 recommends that PWHT should be avoided. The ASME codes, however, specify a PWHT of 552°C to 585°C for both 9%Ni and 5%Ni alloys when thickness exceeds 51mm (2 inches). There are also differences in PWHT requirements in the EN and the ASME specifications for the other types of low temperature steels discussed in this article as tabulated below.

Steel typeEN 13445-4 ASME B31.3

Thickness(mm)

Temperaturerange (°C)

Thickness(mm)

Temperaturerange (°C)

FG C/Mn >35 550-600 >19 593-6491.5%Ni >35 530-580 >19 593-6352.5%Ni >35 530-580 >19 593-6353.5%Ni >35 530-580 >19 593-6355%Ni >35 530-580 51 552-5859%Ni all none 51 552-585Close control of the PWHT temperature is most important as nickel reduces the lower transformation temperature.

Exceeding the specified temperatures, particularly of the 3.5%Ni and above alloys, may cause the parent metal to transform, resulting in a substantial loss of tensile strength.

One significant problem that is frequently encountered with the nickel steels is that of residual magnetism causing arc blow. This is a particular problem with the 9%Ni steel which can become easily and very strongly magnetised, making it impossible to weld with the arc welding processes. Extreme care needs to be taken during handling, transportation and erection to minimise the effect. Use of alternating current during welding can help overcome some of the difficulties but it may be necessary to degauss the area surrounding the weld.

Weld defects/imperfections - incomplete root fusion and penetration

Job Knowledge

The SS Schenectady, an all welded tanker, broke in two whilst lying in dock in 1943. Principal causes of this failure were poor design and bad workmanship

The characteristic features and principal causes of incomplete root fusion and penetration are described. General guidelines on 'best practice' are given so welders can minimise the risk of introducing imperfections during fabrication.

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Fabrication and service defects and imperfections

As the presence of imperfections in a welded joint may not render the component defective in the sense of being unsuitable for the intended application, the preferred term is imperfection rather than defect. For this reason, production quality for a component is defined in terms of a quality level in which the limits for the imperfections are clearly defined, for example Level B, C or D in accordance with the requirements of BS EN ISO 5817. For the American standards ASME IX and AWS D1.1, the acceptance levels are contained in the standards.

The application code will specify the quality levels which must be achieved for the various joints.

Imperfections can be broadly classified into those produced on fabrication of the component or structure and those formed as result of adverse conditions during service. The principal types of imperfections are:

fabrication:

lack of fusion lack of or incomplete penetration cracks porosity inclusions incorrect weld shape and sizeservice:

brittle fracture stress corrosion cracking fatigue failureWelding procedure, joint features and access and welder technique will have a direct effect on fabrication imperfections. Incorrect procedure or poor technique may produce imperfections leading to premature failure in service.

Incomplete root fusion or penetration

Identification

Incomplete root fusion is when the weld fails to fuse one side of the joint in the root. Incomplete root penetration occurs when both sides root region of the joint are unfused. Typical imperfections can arise in the following situations:

an excessively thick root face in a butt weld (Fig. 1a) too small a root gap (Fig. 1b) misplaced welds (Fig. 1c) failure to remove sufficient metal in cutting back to sound metal in a double sided weld (Fig. 1d) incomplete root fusion when using too low an arc energy (heat) input (Fig. 1e) too small a bevel angle, too large a diameter electrode in MMA welding (Fig 2)

a) Excessively thick root face;

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c) Misplaced welds;

b) Too small a root gap;

d) Power input too low;

e) Arc (heat) input too low

Fig. 1 Causes of incomplete root fusion

a) Large diameter electrode;

b) Small diameter electrod

Fig. 2 Effect of electrode diameter on root fusion and

Causes

These types of imperfection are more likely in consumable electrode processes (MIG, MAG, FCAW, MMA and SAW) where the weld metal is 'automatically' deposited as the arc consumes the electrode wire or rod. The welder has limited control of weld pool penetration independent of depositing weld metal. Thus, the non consumable electrode TIG process in which the welder controls the amount of filler material depoisted independent of penetration is less prone to this type of defect.

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In MMA welding, the risk of incomplete root fusion and root penetration can be reduced by using the correct welding parameters and electrode diameter to give adequate arc energy input and satisfactory penetration. Electrode diameter is also important in that it should be small enough to give adequate access to the root, especially when using a small included vee angle (Fig 2). It is common practice to use either a 2.5mm or 3.25mm diameter electrode for the root run so the welder can manipulate the weld pool and control the degree of penetration. However, for the fill passes where penetration requirements are less critical, a 4mm or 5mm diameter electrode may be used to achieve higher deposition rates.

In MIG welding, the correct welding parameters for the material thickness, and a short arc length, should give adequate weld bead penetration. Too low a current level for the size of root face will give inadequate weld penetration. Too high a level, causing the welder to move too quickly, will result in the weld pool bridging the root without achieving adequate penetration.

It is also essential that the correct root face size and bevel angles are used and that the joint root gap is set accurately. To prevent the root gap from closing, adequate tacking will be required.

Best practice in prevention

The following techniques can be used to prevent lack of root fusion:

In TIG welding, do not use too large a root face or too small a root gap and ensure the welding current is sufficient for the weld pool to penetrate fully the root

In MMA welding, use the correct current level and not too large an electrode diameter for the root run

In MIG / MAG welding, use a sufficiently high welding current level which is supported by the appropriate arc voltage for the application 

When using a joint configuration with a root gap, make sure it is of adequate width and does not close up during tacking and subsequent welding

Do not use too low a current level causing the weld pool to bridge the root gap without fully penetrating the root.

Acceptance standards

The limits for lack of or incomplete penetration are specified in BS EN ISO 5817 for the three quality levels.

Lack of or incomplete root penetration is not permitted for Quality Level B (stringent) and Level C (intermediate). However, Level C makes an exception for partial penetration butt welds welded from both sideas.

For Quality Level D (moderate) short lack of or incomplete penetration imperfections are permitted.

Incomplete root penetration is not permitted in the manufacture of pressure vessels but is allowable in the manufacture of pipework depending on material and wall thickness.

Remedial actions

If the root cannot be directly inspected, for example using a penetrant or magnetic particle inspection technique, detection is by radiography or ultrasonic inspection. Remedial action will normally require removal by gouging or grinding to sound metal, followed by re-welding usually in conformity with the original welding procedure.

Weld defects/imperfections in welds - lack of sidewall and inter-run fusion

Job Knowledge

117

Demagnetising a pipe

This article describes the characteristic features and principal causes of lack of sidewall and inter-run fusion. General guidelines on best practice are given so that welders can minimise the risk of imperfections during fabrication.

Identification

Lack of fusion imperfections can occur when the weld metal fails

to fuse completely with the sidewall of the joint (Fig. 1) to penetrate adequately the previous weld bead (Fig. 2).

Fig. 1. Lack of side wall fusion

Fig. 2. Lack of inter-run fusion

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Causes

The principal causes are too narrow a joint preparation, incorrect welding parameter settings, poor welder technique and magnetic arc blow. Insufficient cleaning of oily or scaled surfaces can also contribute to lack of fusion. These types of imperfection are more likely to happen when access to the joint is restricted.

Joint preparation

Too narrow a joint preparation often causes the arc to be attracted to one of the side walls causing lack of side wall fusion on the other side of the joint or inadequate penetration into the previously deposited weld bead. Too great an arc length may also increase the risk of preferential melting along one side of the joint and cause shallow penetration. In addition, a narrow joint preparation may prevent adequate access into the joint or encourage flooding the joint with moulting weld metal. For example, this happens in MMA welding when using a large diameter electrode, or in MIG, MAG and FCAW welding where an allowance has not been made for the diameter of the sheilding gas nozzle. Consideration should also be given to fabrication features that may obstruct the welding torch.

Welding parameters

It is important to use a sufficiently high current for the arc to penetrate into the joint sidewall and previously deposited weld runs. Consequently, too high a welding speed for the welding current will increase the risk of these imperfections. However, too high a current or too low a welding speed will cause weld pool flooding ahead of the arc resulting in poor or non-uniform penetration.

Welder technique

Poor welder technique such as incorrect angle or manipulation of the electrode/welding gun, will prevent adequate fusion of the joint sidewall. Weaving, especially dwelling at the joint sidewall, will enable the weld pool to wash into the parent metal, greatly improving sidewall fusion. It should be noted that the amount of weaving may be restricted by the welding procedure specification limiting the arc energy input, particularly when welding alloy or high notch toughness steels.

Magnetic arc blow

When welding ferromagnetic steels lack of fusion imperfections can be caused through uncontrolled deflection of the arc, usually termed arc blow. Arc deflection can be caused by distortion of the magnetic field produced by the arc current (Fig. 3), through:

residual magnetism in the material through using magnets for handling earth's magnetic field, for example in pipeline welding position of the current return cable clampThe effect of welding past the current return cable which is bolted to the centre of the place is shown in Fig. 4. The interaction of the magnetic field surrounding the arc and that generated by the current flow in the plate to the current return cable is sufficient to deflect the weld bead. Distortion of the arc current magnetic field can be minimised by positioning the current return cable clamp so that welding is always towards or away from the clamp and, in MMA welding, by using AC instead of DC. Often the only effective means is to demagnetise the steel before welding.

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Fig. 3. Interaction of magnetic forces causing arc deflection

Fig. 4. Weld bead deflection in DC MMA welding caused by welding past the current return connection

Best practice in prevention

The following fabrication techniques can be used to prevent formation of lack of sidewall and interrun fusion imperfections:

use a sufficiently wide joint preparation select welding parameters (high current level, short arc length, not too high a welding speed) to

promote penetration into the joint side wall and previousl deposited weld runs without causing flooding ensure the electrode/gun angle and manipulation technique will give adequate side wall fusion use weaving and dwell to improve side wall fusion providing there are no heat input restrictions if arc blow occurs, reposition the current return cable clamp, use AC (in MMA welding) or

demagnetise the steelAcceptance standards

The limits for incomplete fusion imperfections in arc welded joints in steel are specified in BS EN ISO 5817 for the three quality levels (see Table). These types of imperfection are not permitted for Quality Level B (stringent) and C (intermediate). For Quality level D (moderate) they are only permitted providing they are intermittent and not surface breaking.

For arc welded joints in aluminium, long imperfections are not permitted for all three quality levels. However, for quality levels C and D, short imperfections are permitted but the total length of the

imperfections is limited depending on the butt weld or the fillet weld throat thickness. 

Acceptance limits for specific codes and application standards

Application Code/Standard Acceptance limit

Steel BS EN ISO 5817:2007Level B and C not permitted.Level D short imperfections permitted but not surface breaking.

Aluminium BS EN ISO 10042:2005

Levels B, C, D.Long imperfections not permitted.Levels C and D.Short imperfections permitted.

Pressure vessels

BS PD5500:2012+A1: 2012

Not permitted

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Storage tanks BS EN 14015:2004 Not permitted

Pipework BS2633:1994'l' not greater than 15mm(depending on wall thickness)

Line pipe API 1104 (R2010)'l' not greater than 25mm(less when weld length <300mm)

Detection and remedial action

If the imperfections are surface breaking, they can be detected using a penetrant or magnetic particle inspection technique. For sub-surface imperfections, detection is by radiography or ultrasonic inspection. Ultrasonic inspection is normally more effective than radiography in detecting lack of inter-run fusion imperfections.

Remedial action will normally require their removal by localised gouging, or grinding, followed by re-welding as specified in the agreed welding procedure.

If lack of fusion is a persistent problem, and is not caused by magnetic arc blow, the welding procedures should be amended or the welders retrained.

Defects/imperfections in welds - porosity

Job Knowledge

The characteristic features and principal causes of porosity imperfections are described. Best practice guidelines are given so welders can minimise porosity risk during fabrication.

Identification

Porosity is the presence of cavities in the weld metal caused by the freezing in of gas released from the weld pool as it solidifies. The porosity can take several forms: 

distributed surface breaking pores wormhole crater pipesCause and prevention

Distributed porosity and surface pores

Distributed porosity (Fig. 1) is normally found as fine pores throughout the weld bead. Surface breaking pores (Fig. 2)usually indicate a large amount of distributed porosity 

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Fig. 1. Uniformly distributed porosity

Fig. 2. Surface breaking pores (T fillet weld in primed plate)

Cause

Porosity is caused by the absorption of nitrogen, oxygen and hydrogen in the molten weld pool which is then released on solidification to become trapped in the weld metal.

Nitrogen and oxygen absorption in the weld pool usually originates from poor gas shielding. As little as 1% air entrainment in the shielding gas will cause distributed porosity and greater than 1.5% results in gross surface breaking pores. Leaks in the gas line, too high a gas flow rate, draughts and excessive turbulence in the weld pool are frequent causes of porosity.

Hydrogen can originate from a number of sources including moisture from inadequately dried electrodes, fluxes or the workpiece surface. Grease and oil on the surface of the workpiece or filler wire are also common sources of hydrogen.

Surface coatings like primer paints and surface treatments such as zinc coatings, may generate copious amounts of fume during welding. The risk of trapping the evolved gas will be greater in T joints than butt joints especially when fillet welding on both sides (see Fig 2). Special mention should be made of the so-called weldable (low zinc) primers. It should not be necessary to remove the primers but if the primer thickness exceeds the manufacturer's recommendation, porosity is likely to result especially when using welding processes other than MMA.

Prevention

The gas source should be identified and removed as follows: 

Air entrainment

- seal any air leak

- avoid weld pool turbulence

- use filler with adequate level of deoxidants

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- reduce excessively high gas flow

- avoid draughts

Hydrogen

- dry the electrode and flux

- clean and degrease the workpiece surface

Surface coatings

- clean the joint edges immediately before welding

- check that the weldable primer is below the recommended maximum thickness

Elongated pores or wormholes

Wormholes

Characteristically, wormholes are elongated pores (Fig. 3) which produce a herring bone appearance on the radiograph.

Cause

Wormholes are indicative of a large amount of gas being formed which is then trapped in the solidifying weld metal. Excessive gas will be formed from gross surface contamination or very thick paint or primer coatings. Entrapment is more likely in crevices such as the gap beneath the vertical member of a horizontal-vertical, T joint which is fillet welded on both sides.

When welding T joints in primed plates it is essential that the coating thickness on the edge of the vertical member is not above the manufacturer's recommended maximum, typically 20µm, through over-spraying.

Prevention

Eliminating the gas and cavities prevents wormholes.

Gas generation

- clean the workpiece surfaces at and adjacent to the location where the weld will be made

- remove any surface contamination, in particular oil, grease, rust and residue from NDT operations 

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- remove any surface coatings from the joint area to expose bright material 

- check the primer thickness is below the manufacturer's maximum

Joint geometry

- avoid a joint geometry which creates a cavity

 

Crater pipe

A crater pipe forms during the final solidification of the weld pool and is often associated with some gas porosity.

Cause

This imperfection results from shrinkage on weld pool solidification. Consequently, conditions which exaggerate the liquid to solid volume change will promote its formation. Extinquishing the welding arc will result in the rapid solidification of the weld pool.

In TIG welding, autogenous techniques, or stopping the welding wire entering the weld pool before extinquishing the welding arc, will effect crater formation and may promote the pipe imperfection.

Prevention

Crater pipe imperfection can be prevented by controlling the rate at which the welding arc is extinquished  or by welder technique manipulating the welding arc and welding wire

Removal of stop

- use run-off tag to enable the welding arc to be extinquisehd outside the welded joint

- grind out the weld run stop crater before continuing with the next electrode or depositing the subsequent weld run

Welder technique

- progressively reduce the welding current to reduce the weld pool size (use slope-down or crater fill functions) 

- add filler (TIG) to compensate for the weld pool shrinkage 

Porosity susceptibility of materials

Gases likely to cause porosity in the commonly used range of materials are listed in the Table.

Principal gases causing porosity and recommended cleaning methods

Material Gas Cleaning

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C-Mn steel Hydrogen, Nitrogen and Oxygen Grind to remove scale coatings

Stainless steel Hydrogen Degrease + wire brush + degrease

Aluminium and alloys Hydrogen Chemical clean + wire brush + degrease + scrape

Copper and alloys Hydrogen, Nitrogen Degrease + wire brush + degrease

Nickel and alloys Nitrogen Degrease + wire brush + degrease

Detection and remedial action

If the imperfections are surface breaking, they can be detected using a penetrant or magnetic particle inspection technique. For sub surface imperfections, detection is by radiography or ultrasonic inspection. Radiography is normally more effective in detecting and characterising porosity imperfections. However, detection of small pores is difficult especially in thick sections.

Remedial action normally needs removal by localised gouging or grinding but if the porosity is widespread, the entire weld should be removed. The joint should be re-prepared and re-welded as specified in the agreed welding procedure.

Defects/imperfections in welds - slag inclusions

Job Knowledge

Prevention of slag inclusions by grinding between runs

The characteristic features and principal causes of slag imperfections are described.

Identification

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Fig. 1. Radiograph of a butt weld showing two slag lines in the weld root

Slag is normally seen as elongated lines either continuous or discontinuous along the length of the weld. This is readily identified in a radiograph, Fig 1. Slag inclusions are usually associated with the flux processes, ie MMA, FCA and submerged arc, but they can also occur in MIG welding.

Causes

As slag is the residue of the flux coating in MMA welding, it is principally a deoxidation product from the reaction between the flux, air and surface oxide. The slag becomes trapped in the weld when two adjacent weld beads are deposited with inadequate overlap and a void is formed. When the next layer is deposited, the entrapped slag is not melted out. Slag may also become entrapped in cavities in multi-pass welds through excessive undercut in the weld toe or the uneven surface profile of the preceding weld runs, Fig 2.

As they both have an effect on the ease of slag removal, the risk of slag imperfections is influenced by

Type of flux coating Welder techniqueThe type and configuration of the joint, welding position and access restrictions all have an influence on the risk of slag imperfections.

Fig. 2. The influence of welder technique on the risk of slag inclusions when welding with a basic MMA (E7018) electrode a) Poor (convex) weld bead profile resulted in pockets of slag being trapped between the

weld runs

b) Smooth weld bead profile allows the slag to be readily removed between runs

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Type of flux coating

One of the main functions of the flux coating in welding is to produce a slag which will flow freely over the surface of the weld pool to protect it from oxidation. As the slag affects the handling characteristics of the MMA electrode, its surface tension and freezing rate can be equally important properties. For welding in the flat and horizontal/vertical positions, a relatively viscous slag is preferred as it will produce a smooth weld bead profile, is less likely to be trapped and, on solidifying, is normally more easily removed. For vertical welding, the slag must be more fluid to flow out to the weld pool surface but have a higher surface tension to provide support to the weldpool and be fast freezing.

The composition of the flux coating also plays an important role in the risk of slag inclusions through its effect on the weld bead shape and the ease with which the slag can be removed. A weld pool with low oxygen content will have a high surface tension producing a convex weld bead with poor parent metal wetting. Thus, an oxidising flux, containing for example iron oxide, produces a low surface tension weld pool with a more concave weld bead profile, and promotes wetting into the parent metal. High silicate flux produces a glass-like slag, often self detaching. Fluxes with a lime content produce an adherent slag which is difficult to remove.

The ease of slag removal for the principal flux types are:

Rutile or acid fluxes - large amounts of titanium oxide (rutile) with some silicates. The oxygen level of the weld pool is high enough to give flat or slightly convex weld bead. The fluidity of the slag is determined by the calcium fluoride content. Fluoride-free coatings designed for welding in the flat position produce smooth bead profiles and an easily removed slag. The more fluid fluoride slag designed for positional welding is less easily removed.

Basic fluxes - the high proportion of calcium carbonate (limestone) and calcium fluoride (fluospar) in the flux reduces the oxygen content of the weld pool and therefore its surface tension. The slag is more fluid than that produced with the rutile coating. Fast freezing also assists welding in the vertical and overhead positions but the slag coating is more difficult to remove.

Consequently, the risk of slag inclusions is significantly greater with basic fluxes due to the inherent convex weld bead profile and the difficulty in removing the slag from the weld toes especially in multi-pass welds.

Welder technique

Welding technique has an important role to play in preventing slag inclusions. Electrode manipulation should ensure adequate shape and degree of overlap of the weld beads to avoid forming pockets which can trap the slag. Thus, the correct size of electrode for the joint preparation, the correct angle to the workpiece for good penetration and a smooth weld bead profile are all essential to prevent slag entrainment.

In multi-pass vertical welding, especially with basic electrodes, care must be taken to fuse out any remaining minor slag pockets and minimise undercut. When using a weave, a slight dwell at the extreme edges of the weave will assist sidewall fusion and produce a flatter weld bead profile.

Too high a current together with a high welding speed will also cause sidewall undercutting which makes slag removal difficult.

It is crucial to remove all slag before depositing the next run. This can be done between runs by grinding, light chipping or wire brushing. Cleaning tools must be identified for different materials eg steels or stainless steels, and segregated.

When welding with difficult electrodes, in narrow vee butt joints or when the slag is trapped through undercutting, it may be necessary to grind the surface of the weld between layers to ensure complete slag removal.

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Best practice

The following techniques can be used to prevent slag inclusions:

Use welding techniques to produce smooth weld beads and adequate inter-run fusion to avoid forming pockets to trap the slag

Use the correct current and travel speed to avoid undercutting the sidewall which will make the slag difficult to remove

Remove slag between runs paying particular attention to removing any slag trapped in crevices Use grinding when welding difficult butt joints otherwise wire brushing or light chipping may be

sufficient to remove the slag.Acceptance standards

Slag and flux inclusions are linear defects but because they do not have sharp edges compared with cracks, they may be permitted by specific standards and codes. The limits in steel are specified in BE EN ISO 5817: 2007 for three quality levels.

Defects - solidification cracking

Job Knowledge

Weld repair on a cast iron exhaust manifold

A crack may be defined as a local discontinuity produced by a fracture which can arise from the stresses generated on cooling or acting on the structure. It is the most serious type of imperfection found in a weld and should be removed. Cracks not only reduce the strength of the weld through the reduction in the cross section thickness but also can readily propagate through stress concentration at the tip, especially under impact loading or during service at low temperature.

Identification

Visual appearance

Solidification cracks are normally readily distinguished from other types of cracks due to the following characteristic factors:

they occur only in the weld metal they normally appear as straight lines along the centreline of the weld bead, as shown in Fig.1, but

may occasionally appear as transverse cracking depending on the solidification structure solidification cracks in the final crater may have a branching appearance as the cracks are often 'open', they can be visible to the naked eye

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Fig.1 Solidification crack along the centre line of the weld

On breaking open the weld, the crack surface in steel and nickel alloys may have a blue oxidised appearance, showing that they were formed while the weld metal was still hot.

Metallography

The cracks form at the solidification boundaries and are characteristically interdendritic. The morphology reflects the weld solidification structure and there may be evidence of segregation associated with the solidification boundary.

Causes

The overriding cause of solidification cracking is that the weld bead in the final stage of solidification has insufficient strength to withstand the contraction stresses generated as the weld pool solidifies. Factors which increase the risk include:

insufficient weld bead size or shape welding under high restraint material properties such as a high impurity content or a relatively large amount of shrinkage on

solidification.Joint design can have a significant influence on the level of residual stresses. Large gaps between component parts will increase the strain on the solidifying weld metal, especially if the depth of penetration is small. Therefore, weld beads with a small depth-to-width ratio, such as formed in bridging a large gap with a wide, thin bead, will be more susceptible to solidification cracking, as shown in Fig.2. In this case, the centre of the weld which is the last part to solidify, is a narrow zone with negligible cracking resistance.

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Fig.2 Weld bead penetration too small

Segregation of impurities to the centre of the weld also encourages cracking. Concentration of impurities ahead of the solidifying weld front forms a liquid film of low freezing point which, on solidification, produces a weak zone. As solidification proceeds, the zone is likely to crack as the stresses through normal thermal contraction build up. If liquid from the weld pool can feed into an incipent crack, it can be prevented. For this reason, an elliptically shaped weld pool is preferable to a tear drop shape, and fast welding speeds, which result in a large separation between the weld pool and cracking locations, increase the risk of cracking. Welding with contaminants such as cutting oils on the surface of the parent metal will also increase the build up of impurities in the weld pool and the risk of cracking. 

As the compositions of the plate and the filler determine the weld metal composition they will, therefore, have a substantial influence on the susceptibility of the material to cracking.

Steels

Cracking is associated with impurities, particularly sulphur and phosphorus, and is promoted by carbon whereas manganese and silicon can help to reduce the risk. To minimise the risk of cracking, fillers with low carbon and impurity levels and a relatively high manganese content are preferred. As a general rule, for carbon-manganese steels, the total sulphur and phosphorus content should be no greater than 0.06%.

Weld metal composition is dominated by the consumable and as the filler is normally cleaner than the metal being welded, cracking is less likely with low dilution processes such as MMA and MIG. Plate composition assumes greater importance in high dilution situations such as when welding the root in butt welds, using an autogenous welding technique like TIG, or a high dilution process such as submerged arc welding.

In submerged arc welds, as described in EN 1011-2:2001 Annex E, the cracking risk may be assessed by calculating the Units of Crack Susceptibility (UCS) from the weld metal chemical composition (weight %):

UCS = 230C* + 190S + 75P + 45Nb - 12.3Si - 5.4Mn - 1C* = carbon content or 0.08 whichever is higher

Although arbitrary units, a value of <10 indicates high cracking resistance whereas >30 indicates a low resistance. Within this range, the risk will be higher in a weld run with a high depth to width ratio, made at high welding speeds or where the fit-up is poor. For fillet welds, runs having a depth to width ratio of about one, UCS values of 20 and above will indicate a risk of cracking. For a butt weld, values of about 25 UCS are critical. If the depth to width ratio is decreased from 1 to 0.8, the allowable UCS is increased by about nine. However, very low depth to width ratios, such as obtained when penetration into the root is not achieved, also promote cracking.

Aluminium

The high thermal expansion (approximately twice that of steel) and substantial contraction on solidification (typically 5% more than in an equivalent steel weld) means that aluminium alloys are more prone to cracking. The risk can be reduced by using a crack resistant filler (usually from the 4xxx and 5xxx series alloys) but the disadvantage is that the resulting weld metal is likely to have non-matching properties such as a lower strength than the parent metal.

Austenitic Stainless Steel

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A fully austenitic stainless steel weld is more prone to cracking than one containing between 5-10% of ferrite. The beneficial effect of ferrite has been attributed to its capacity to contain harmful impurities within the grains which would otherwise form low melting point segregates and consequently interdendritic cracks. Therefore the choice of filler material is important to suppress cracking so a type 308 filler is used to weld type 304 stainless steel.

Best practice in avoiding solidification cracking

Apart from the choice of material and filler, the principal techniques for minimising the risk of welding solidification cracking are:

Control joint fit-up to reduce gaps. Before welding, clean off all contaminants from the material Ensure that the welding sequence will not lead to a build-up of thermally induced stresses. Select welding parameters and technique to produce a weld bead with an adequate depth to width

ratio, or with sufficient throat thickness (fillet weld), to ensure the weld bead has sufficient resistance to the solidification stresses (recommend a depth to width ratio of at least 0.5:1).

Avoid producing too large a depth to width ratio which will encourage segregation and excessive transverse strains in restrained joints. As a general rule, weld beads whose depth to width ratio exceeds 2:1 will be prone to solidification cracking.

Avoid high welding speeds (at high current levels) which increase the amount of segregation and the stress level across the weld bead.

At the run stop, ensure adequate filling of the crater to avoid an unfavourable concave shape.Acceptance standards

As solidification cracks and crater cracks are linear imperfections with sharp edges, they are not permitted for welds meeting the quality levels B, C and D in accordance with the requirements of BS EN ISO 5817:2007. Crater pipes may be permitted for quality level D, depending on their size.

Detection and remedial action

Surface breaking solidification cracks can be readily detected using visual examination, liquid penetrant or magnetic particle testing techniques. Internal cracks require ultrasonic or radiographic examination techniques.

Most codes will specify that all cracks should be removed. A cracked component should be repaired by removing the cracks with a safety margin of approximately 5mm beyond the visible ends of the crack. The excavation is then re-welded using a filler which will not produce a crack sensitive deposit.

Defects - hydrogen cracks in steels - identification

Job Knowledge

131

Preheating to avoid hydrogen cracking

Hydrogen cracking may also be called cold cracking or delayed cracking. The principal distinguishing feature of this type of crack is that it occurs in ferritic steels, most often immediately on welding or a short time after welding.

In this issue, the characteristic features and principal causes of hydrogen cracks are described.

Identification

Visual appearance

Hydrogen cracks can be usually be distinguished due to the following characteristics:

In C-Mn steels, the crack will normally originate in the heat affected zone (HAZ), but may extend into the weld metal(Fig 1).

Cracks can also occur in the weld bead, normally transverse to the welding direction at an angle of 45° to the weld surface. They follow a jagged path, but may be non-branching.

In low alloy steels, the cracks can be transverse to the weld, perpendicular to the weld surface, but are non-branching, and essentially planar.

Fig. 1 Hydrogen cracks originating in the HAZ and weld metal. (Note that the type of cracks shown would not be expected to form in the same weldment.)

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On breaking open the weld (prior to any heat treatment), the surface of the cracks will normally not be oxidised, even if they are surface breaking, indicating they were formed when the weld was at or near ambient temperature. A slight blue tinge may be seen from the effects of preheating or welding heat.

Metallography

Cracks which originate in the HAZ are usually associated with the coarse grain region, (Fig 2). The cracks can be intergranular, transgranular or a mixture. Intergranular cracks are more likely to occur in the harder HAZ structures formed in low alloy and high carbon steels. Transgranular cracking is more often found in C-Mn steel structures.

In fillet welds, cracks in the HAZ are usually associated with the weld root and parallel to the weld. In butt welds, the HAZ cracks are normally oriented parallel to the weld bead.

Fig. 2 Crack along the coarse grain structure in the HAZ

Causes

There are three factors which combine to cause cracking:

hydrogen generated by the welding process a hard brittle structure which is susceptible to cracking tensile stresses acting on the welded jointCracking usually occurs at temperatures at or near normal ambient. It is caused by the diffusion of hydrogen to the highly stressed, hardened part of the weldment.

In C-Mn steels, because there is a greater risk of forming a brittle microstructure in the HAZ, most of the hydrogen cracks are to be found in the parent metal. With the correct choice of electrodes, the weld metal will have a lower carbon content than the parent metal and, hence, a lower carbon equivalent (CE). However, transverse weld metal cracks can occur, especially when welding thick section components; the risk of cracking is increased if the weld metal carbon content exceeds that of the parent steel.

In low alloy steels, as the weld metal structure is more susceptible than the HAZ, cracking may be found in the weld bead.

The main factors which influence the risk of cracking are:

weld metal hydrogen parent material composition parent material thickness

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stresses acting on the weld during welding or imposed (shortly) after welding heat inputWeld metal hydrogen content

The principal source of hydrogen is moisture contained in the flux, i.e. the coating of MMA electrodes, the flux in cored wires and the flux used in submerged arc welding. The amount of hydrogen generated is influenced by the electrode type. Basic electrodes normally generate less hydrogen than rutile and cellulosic electrodes.

It is important to note that there can be other significant sources of hydrogen, e.g. from the material, where processing or service history has left the steel with a significant level of hydrogen or moisture from the atmosphere. Hydrogen may also be derived from the surface of the material or the consumable.

Sources of hydrogen will include:

oil, grease and dirt rust paint and coatings cleaning fluidsParent metal composition

This will have a major influence on hardenability and, with high cooling rates, the risk of forming a hard brittle structure in the HAZ. The hardenability of a material is usually expressed in terms of its carbon content or, when other elements are taken into account, its carbon equivalent (CE) value.

The higher the CE value, the greater the risk of hydrogen cracking. Generally, steels with a CE value of <0.4 are not susceptible to HAZ hydrogen cracking, as long as low hydrogen welding consumables or processes are used.

Parent material thickness

Material thickness will influence the cooling rate and therefore the hardness level, the microstructure produced in the HAZ and the level of hydrogen retained in the weld.

The 'combined thickness' of the joint, ie the sum of the thicknesses of material meeting at the joint line, will determine, together with the joint geometry, the cooling rate of the HAZ and its hardness. Consequently, as shown inFig. 3, a fillet weld is likely to have a greater risk than a butt weld in the same material thickness.

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Fig.3 Combined thickness measurements for butt and fillet joints

Stresses acting on the weld

Cracks are more likely to initiate at regions of stress concentration, particularly at the toe and root of the weld.

The stresses generated across the welded joint as it contracts will be greatly influenced by external restraint, material thickness, joint geometry and fit-up. Poor fit-up (excessive root gap) in fillet welds markedly increases the risk of cracking. The degree of restraint acting on a joint will generally increase as welding progresses, due to the increase in stiffness of the fabrication.

Heat input

The heat input to the material from the welding process, together with the material thickness and preheat temperature, will determine the thermal cycle and the resulting microstructure and hardness of both the HAZ and the weld metal.

Increasing the heat input will reduce the hardness level, and therefore reduce the risk of HAZ cracking. However, as the diffusion distance for the escape of hydrogen from a weld bead increases with increasing heat input, the risk of weld metal cracking is increased.

Heat input per unit length is calculated by multiplying the arc energy by a thermal efficiency factor, according to the following formula: 

V = arc voltage (V)A = welding current (A)S = welding speed (mm/min)k = thermal efficiency factor

In calculating heat input, the thermal efficiency must be taken into consideration. The thermal efficiency factors given in EN 1011-1: 2009 for the principal arc welding processes, are:

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Submerged arc(single wire)

1.0

MMA 0.8

MIG/MAG and flux cored wire 0.8

TIG and plasma 0.6

In MMA welding, heat input is normally controlled by means of the run-out length from each electrode, which is proportional to the heat input. As the run-out length is the length of weld deposited from one electrode, it will depend upon the welding technique, e.g. weave width /dwell.

Defects - hydrogen cracks in steels - prevention and best practice

Job Knowledge

Preheating of a jacket structure to prevent hydrogen cracking

Techniques and practical guidance on the avoidance of hydrogen cracks are described.

Preheating, interpass and post heating to prevent hydrogen cracking

There are three factors which combine to cause hydrogen cracking in arc welding:

hydrogen generated by the welding process a hard brittle structure which is susceptible to cracking tensile stresses acting on the welded jointCracking generally occurs when the temperature has reached normal ambient. In practice, for a given situation (material composition, material thickness, joint type, electrode composition and heat input), the risk of hydrogen cracking is reduced by heating the joint.

Preheat

Preheat, which slows the cooling rate, allows some hydrogen to diffuse away, and generally reduces the hardness, and therefore susceptibility to cracking, of hard, crack-sensitive microstructural regions. The recommended levels of preheat for carbon and carbon manganese steel are detailed in EN 1011-2: 2001

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(which incorporates nomograms derived from those in BS 5135: 1984). The preheat level may be as high as 200°C for example, when welding thick section steels with a high carbon equivalent (IIW CE) value.

Interpass and post-heating

As cracking rarely occurs at temperatures above ambient, maintaining the temperature of the weldment during fabrication is equally important. For susceptible steels, it is usually appropriate to maintain the preheat temperature for a given period, typically between two to three hours, to enable the hydrogen to diffuse away from the weld area. In crack-sensitive situations, such as welding higher IIW CE steels or under high restraint conditions, the temperature and heating period should be increased, typically 250-300°C for three to four hours.

For many steels, post-weld heat treatment (PWHT) may be used immediately on completion of welding, i.e. without allowing the preheat temperature to fall. However, in practice, as inspection can only be carried out at ambient temperature, there is the risk that 'rejectable' defects will only be found after PWHT. Also, for highly hardenable steels, a second heat treatment may be required to temper the hard microstructure present after the first PWHT.

Under certain conditions, more stringent procedures (with a higher preheat temperature and/or a lower weld metal hydrogen level) are needed to avoid cracking than those derived from the nomograms for estimating preheat in Fig. C2 of EN 1011-2. Section C.2.9 of this standard mentions the following conditions:

1. high restraint, including welds in section thicknesses above approximately 50mm, and root runs in double bevel joints

2. thick sections (≥ approximately 50mm)

3. low carbon equivalent steels (C-Mn steels with C ≤ 0.1% and IIW CE ≤ approximately 0.42)

4. 'clean' or low sulphur steels (S ≤ approximately 0.008%), as a low sulphur and low oxygen content will increase the hardenability of a steel.

5. alloyed weld metal where preheat levels to avoid HAZ cracking may be insufficient to protect the weld metal. Low hydrogen processes and consumables should be used. Schemes for predicting the preheat requirements to avoid weld metal cracking generally require the weld metal diffusible hydrogen level and the weld metal tensile strength as input. 

Use of austenitic and nickel alloy weld metal to prevent cracking

In situations where preheating is impractical, or does not prevent cracking, it will be necessary to use an austenitic consumable. Austenitic stainless steel and nickel alloy electrodes will produce a weld metal which at ambient temperature has a higher solubility for hydrogen than ferritic steel. Thus, any hydrogen formed during welding becomes locked in the weld metal, with very little diffusing to the HAZ on cooling to ambient temperature.

A commonly used austenitic MMA electrode is 23Cr:12Ni, e.g. from EN 1600: 1997. However, as nickel alloys have a lower coefficient of thermal expansion than stainless steel, nickel alloy electrodes are preferred, to reduce the shrinkage strain, when welding highly restrained joints. Figure 1 is a general guide on the levels of preheat when using austenitic electrodes. When welding steels with up to 0.2%C, a preheat would not normally be required. However, above 0.4%C a minimum temperature of 150°C will be needed to prevent HAZ cracking. The influence of hydrogen level and the degree of restraint are also illustrated in the figure.

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Fig.1 Guide to preheat temperature when using austenitic MMA electrodes at 1-2kJ/mm a) low restraint (e.g. material thickness <30mm) b) high restraint (e.g. material thickness >30mm)

Best practice in avoiding hydrogen cracking

Reduction in weld metal hydrogen

The most effective means of avoiding hydrogen cracking is to reduce the amount of hydrogen generated by the consumable, ie by using a low hydrogen process or low hydrogen electrodes.

Welding processes can be classified as high, medium, low, very low and ultra low, depending on the amount of weld metal hydrogen produced in a standard test block. The weld metal diffusible hydrogen levels (ml/100g of deposited metal, measured in a test weld, as specified in BS EN ISO 3690:2001), and the hydrogen scale designations of EN 1011-2: 2001 are as follows:

High >15 Scale A

Medium >10 <15 Scale B

Low >5 <10 Scale C

Very low >3 <5 Scale D

Ultra-low ≤3 Scale E

Figure 2 from Bailey et al illustrates the relative amounts of weld metal hydrogen produced by the major welding processes. MMA, in particular, has the potential to generate a wide range of hydrogen levels. Thus, to achieve the lower values, it is essential that basic electrodes are used, and they are baked in accordance with the manufacturer's recommendations, or taken from special packaging immediately before use, and exposed to ambient conditions for no longer than the time period specified by the manufacturer. For the MIG process, cleaner wires will be required to achieve very low hydrogen levels.

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Fig.2 General relationships between potential hydrogen and weld metal hydrogen levels for arc welding processes

General guidelines

The following general guidelines are recommended for the various types of steel, but requirements for specific steels should be checked according to EN 1011-2: 2001 -

Mild steel (CE <0.4)

- readily weldable, preheat generally not required if low hydrogen processes or electrodes are used

- preheat may be required when welding thick section material, high restraint and with higher levels of hydrogen being generated

C-Mn, medium carbon, low alloy steels (CE 0.4 to 0.5)- thin sections can be welded without preheat, but thicker sections will require low preheat levels, and low hydrogen processes or electrodes should be used

Higher carbon and alloyed steels (CE >0.5)- preheat, low hydrogen processes or electrodes, post-weld heating and slow cooling required.

More detailed guidance on the avoidance of hydrogen cracking is described in EN 1011-2: 2001.

Practical Techniques

The following practical techniques are recommended to avoid hydrogen cracking:

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clean the joint faces and remove contaminants such as paint, cutting oils, grease use a low hydrogen process, if possible bake the electrodes (MMA) or the flux (submerged arc) and then either store them warm or restrict

the duration of exposure to ambient conditions, all in accordance with the manufacturer's recommendations

reduce stresses on the weld by avoiding large root gaps and high restraint if preheating is specified in the welding procedure, it should also be applied when tacking or using

temporary attachments preheat the joint to a distance of at least 75mm from the joint line, ensuring uniform heating through

the thickness of the material measure the preheat temperature on the face opposite that being heated. Where this is impractical,

allow time for the equalisation of temperature after removing the preheating before the temperature is measured

adhere to the preheat and minimum interpass temperature, and heat input requirements maintain heat for approximately two to four hours after welding, depending on crack sensitivity In situations where adequate preheating is impracticable, or cracking cannot be avoided, austenitic

electrodes may be usedAcceptance standards

As hydrogen cracks are linear imperfections which have sharp edges, they are not permitted for welds meeting the quality levels B, C and D in accordance with the requirements of EN ISO 5817.

Detection and remedial action

As hydrogen cracks are often very fine and may be sub-surface, they can be difficult to detect. Surface-breaking hydrogen cracks can be readily detected using visual examination, liquid penetrant or magnetic particle testing techniques. Internal cracks require ultrasonic or radiographic examination techniques. Ultrasonic examination is preferred, as radiography is restricted to detecting relatively wide cracks that are parallel to the beam. As the formation of cracks may be delayed for many hours after completion of welding, the delay time before inspection, according to the relevant fabrication code, should be observed.

Most codes will specify that all cracks should be removed. A cracked component should be repaired by removing the cracks with a safety margin of approximately 5mm beyond the visible ends of the crack. The excavation is then re-welded.

To make sure that cracking does not re-occur, welding should be carried out with the correct procedure, i.e. preheat and an adequate heat input level for the material type and thickness. However, as the level of restraint will be greater and the interpass time shorter when welding within an excavation compared to welding the original joint, it is recommended that a higher level of preheat is used (typically by 50°C).

Defects - lamellar tearing

Job Knowledge

140

BP Forties platform lamellar tears were produced when attempting the repair of lack of root penetration in a brace weld

Lamellar tearing can occur beneath the weld especially in rolled steel plate which has poor through-thickness ductility. The characteristic features, principal causes and best practice in minimising the risk of lamellar tearing are described.

Identification

Visual appearance

The principal distinguishing feature of lamellar tearing is that it occurs in T-butt and fillet welds normally observed in the parent metal parallel to the weld fusion boundary and the plate surface , (Fig 1). The cracks can appear at the toe or root of the weld but are always associated with points of high stress concentration.

Fracture face

The surface of the fracture is fibrous and 'woody' with long parallel sections which are indicative of low parent metal ductility in the through-thickness direction, (Fig 2).

Fig. 1. Lamellar tearing in T butt weld

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Fig. 2. Appearance of fracture face of lamellar tear

Metallography

As lamellar tearing is associated with a high concentration of elongated inclusions oriented parallel to the surface of the plate, tearing will be transgranular with a stepped appearance.

Causes

It is generally recognised that there are three conditions which must be satisfied for lamellar tearing to occur:

1. Transverse strain - the shrinkage strains on welding must act in the short direction of the plate ie through the plate thickness

2. Weld orientation - the fusion boundary will be roughly parallel to the plane of the inclusions

3. Material susceptibility - the plate must have poor ductility in the through-thickness direction

Thus, the risk of lamellar tearing will be greater if the stresses generated on welding act in the through-thickness direction. The risk will also increase the higher the level of weld metal hydrogen

Factors to be considered to reduce the risk of tearing

The choice of material, joint design, welding process, consumables, preheating and buttering can all help reduce the risk of tearing.

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Material

Fig. 3. Relationship between the STRA and sulphur content for 12.5 to 50mm thick plate

Tearing is only encountered in rolled steel plate and not forgings and castings. There is no one grade of steel that is more prone to lamellar tearing but steels with a low Short Transverse Reduction in Area (STRA), commonly associated with a high concentration of rolled sulphide or oxide inclusions, will be susceptible. As a general rule, steels with STRA over 20% are essentially resistant to tearing whereas steels with below 10 to 15% STRA should only be used in lightly restrained joints (Fig. 3).

Steels with a higher strength have a greater risk especially when the thickness is greater than 25mm. Aluminium treated steels with low sulphur contents (<0.005%) will have a low risk.

Steel suppliers can provide plate which has been through-thickness tested with a guaranteed STRA value of over 20%.

Joint Design

Lamellar tearing occurs in joints producing high through-thickness strain, eg T joints or corner joints. In T or cruciform joints, full penetration butt welds will be particularly susceptible. The cruciform structures in which the susceptible plate cannot bend during welding will also greatly increase the risk of tearing.

In butt joints, as the stresses on welding do not act through the thickness of the plate, there is little risk of lamellar tearing.

As angular distortion can increase the strain in the weld root and or toe, tearing may also occur in thick section joints where the bending restraint is high.

Several examples of good practice in the design of welded joints are illustrated in Fig. 4.

As tearing is more likely to occur in full penetration T butt joints, if possible, use two fillet welds, Fig. 4a. 

Double-sided welds are less susceptible than large single-sided welds and balanced welding to reduce the stresses will further reduce the risk of tearing especially in the root, Fig. 4b 

Large single-side fillet welds should be replaced with smaller double-sided fillet welds, Fig. 4c 

Redesigning the joint configuration so that the fusion boundary is more normal to the susceptible plate surface will be particularly effective in reducing the risk, Fig. 4d 

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Fig. 4 Recommended joint configurations to reduce the risk of lamellar tearing - Fig. 4a

Fig. 4c

Fig. 4b

Fig. 4d

Weld size

Lamellar tearing is more likely to occur in large welds typically when the leg length in fillet and T butt joints is greater than 20mm. As restraint will contribute to the problem, thinner section plate which is less susceptible to tearing, may still be at risk in high restraint situations.

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Welding process

As the material and joint design are the primary causes of tearing, the choice of welding process has only a relatively small influence on the risk. However, higher heat input processes which generate lower stresses through the larger HAZ and deeper weld penetration can be beneficial.

As weld metal hydrogen will increase the risk of tearing, a low hydrogen process should be used when welding susceptible steels.

Consumable

Where possible, the choice of a lower strength consumable can often reduce the risk by accommodating more of the strain in the weld metal. A smaller diameter electrode which can be used to produce a smaller leg length, has been used to prevent tearing.

A low hydrogen consumable will reduce the risk by reducing the level of weld metal diffusible hydrogen. The consumables must be dried in accordance with the manufacturer's recommendations.

Preheating

Preheating will have a beneficial effect in reducing the level of weld metal diffusible hydrogen. However, it should be noted that in a restrained joint, excessive preheating could have a detrimental effect by increasing the level of restraint produced by the contraction across the weld on cooling.

Preheating should, therefore, be used to reduce the hydrogen level but it should be applied so that it will not increase the amount of contraction across the weld.

Buttering

Buttering the surface of the susceptible plate with a low strength weld metal has been widely employed. As shown for the example of a T butt weld (Fig. 5) the surface of the plate may be grooved so that the buttered layer will extend 15 to 25mm beyond each weld toe and be about 5 to 10mm thick.

Fig. 5. Buttering with low strength weld metal a) general deposit on the surface of the susceptible plate

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b) in-situ buttering

In-situ buttering ie where the low strength weld metal is deposited first on the susceptible plate before filling the joint, has also been successfully applied. However, before adopting either buttering technique, design calculations should be carried out to ensure that the overall weld strength will be acceptable.

Acceptance standards

As lamellar tears are linear imperfections which have sharp edges, they are not permitted for welds meeting the quality levels B, C and D in accordance with the requirements of BS EN ISO 5817:2007.

Detection and remedial action

If surface-breaking, lamellar tears can be readily detected using visual examination, liquid penetrant or magnetic particle testing techniques. Internal cracks require ultrasonic examination techniques but there may be problems in distinguishing lamellar tears from inclusion bands. The orientation of the tears normally makes them almost impossible to detect by radiography.

Defects/imperfections in welds - reheat cracking

Job Knowledge

Location of reheat cracks in a nuclear pressure vessel steel

The characteristic features and principal causes of reheat cracking are described. General guidelines on best practice are given so that welders can minimise the risk of reheat cracking in welded fabrications.

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Identification

Visual appearance

Reheat cracking may occur in low alloy steels containing alloying additions of chromium and molybdenum or chromium, molybdenum and vanadium when the welded component is being subjected to post weld heat treatment, such as stress relief heat treatment, or has been subjected to high temperature service (typically in the range 350 to 550°C).

Cracking is almost exclusively found in the coarse grained regions of the heat affected zone (HAZ) beneath the weld, or cladding, and in the coarse grained regions within the weld metal. The cracks can often be seen visually, usually associated with areas of stress concentration such as the weld toe.

Cracking may be in the form of coarse macro-cracks or colonies of micro-cracks.

A macro-crack will appear as a 'rough' crack, often with branching, following the coarse grain region, (Fig. 1a). Cracking is always intergranular along the prior austenite grain boundaries (Fig. 1b). Macro-cracks in the weld metal can be oriented either longitudinal or transverse to the direction of welding. Cracks in the HAZ, however, are always parallel to the direction of welding.

Fig.1a. Cracking associated with the coarse grained heat affected zone

Fig.1b. Intergranular morphology of reheat cracks

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Micro-cracking can also be found both in the HAZ and within the weld metal. Micro-cracks in multipass welds will be found associated with the grain coarsened regions which have not been refined by subsequent passes.

Causes

The principal cause is that when heat treating susceptible steels, the grain interior becomes strengthened by carbide precipitation, forcing the relaxation of residual stresses by creep deformation at the grain boundaries.

The presence of impurities which segregate to the grain boundaries and promote temper embrittlement, e.g. antimony, arsenic, tin, sulphur and phosphorus, will increase the susceptibility to reheat cracking.

The joint design can increase the risk of cracking. For example, joints likely to contain stress concentration, such as partial penetration welds, are more liable to initiate cracks.

The welding procedure also has an influence. Large weld beads are undesirable, as they produce coarse columnar grains within the weld metal and a coarse grained HAZ which is less likely to be refined by the subsequent pass, and therefore will be more susceptible to reheat cracking.

Best practice in prevention

The risk of reheat cracking can be reduced through the choice of steel, specifying the maximum impurity level and by adopting a more tolerant welding procedure / technique.

Steel choice

If possible, avoid welding steels known to be susceptible to reheat cracking. For example, A 508 Class 2 is known to be particularly susceptible to reheat cracking, whereas cracking associated with welding and cladding in A508 Class 3 is largely unknown. The two steels have similar mechanical properties, but A508 Class 3 has a lower Cr content and a higher manganese content.

Similarly, in the higher strength, creep-resistant steels, an approximate ranking of their crack susceptibility is as follows:

5 Cr 1Mo lower risk

2.25Cr 1 Mo ↓

0.5Mo B ↓

0.5Cr 0.5Mo 0.25V higher risk

Thus, in selecting a creep-resistant, chromium molybdenum steel, 0.5Cr 0.5Mo 0.25V steel is known to be susceptible to reheat cracking but the 2.25Cr 1Mo which has a similar creep resistance, is significantly less susceptible.

Unfortunately, although some knowledge has been gained on the susceptibility of certain steels, the risk of cracking cannot be reliably predicted from the chemical composition. Various indices, including ΔG1, PSR and Rs, have been used to indicate the susceptibility of steel to reheat cracking. Steels which have a value of ΔG1 of less than 2, PSR less than zero or Rs less than 0.03, are less susceptible to reheat cracking

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ΔG1 = 10C + Cr + 3.3Mo + 8.1V - 2

PSR = Cr +Cu + 2Mo + 10V +7Nb + 5Ti - 2

Rs = 0.12Cu +0.19S +0.10As + P +1.18Sn + 1.49Sb

Irrespective of the steel type, it is important to purchase steels specified to have low levels of impurity elements (antimony, arsenic, tin, bismuth, sulphur and phosphorus). To avoid weld metal reheat cracking, it is necessary to ensure that welding consumables deposit weld metal with appropriately low levels of these impurities, and preferably to avoid coarse columnar grains. Following several instances of weld metal reheat cracking in thick-wall 2.25%Cr-1%Mo-0.25%V reactor vessels, impurities in the flux were identified as being responsible for the cracking, and an equation given for the desired upper limit of these additional impurities.

K = Pb + Bi + 0.03Sb (ppm)

The compositional factor K must be less than 1.5 to achieve freedom from this form of cracking.

Welding procedure and technique

The welding procedure can be used to minimise the risk of reheat cracking by

Producing the maximum refinement of the coarse grain HAZ Limiting the degree of austenite grain growth Eliminating stress concentrationsThe procedure should aim to refine the coarse grained HAZ by subsequent passes. In butt welds, maximum refinement can be achieved by using a steep-sided joint preparation with a low angle of attack to minimise penetration into the side-wall, ( Fig 2a). In comparison, a larger angle V preparation produces a wider HAZ, limiting the amount of refinement achieved by subsequent passes, ( Fig 2b). Narrow joint preparations, however, are more difficult to weld, due to the increased risk of lack of side-wall fusion.

Fig.2a. Welding in the flat position - high degree of HAZ refinement

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Fig.2b. Welding in the horizontal/vertical position - low degree of HAZ refinement

Refinement of the HAZ can be promoted by first buttering the surface of the susceptible plate with a thin weld metal layer using a small diameter (3.2mm) electrode. The joint is then completed using a larger diameter (4 - 4.8mm) electrode, which is intended to generate sufficient heat to refine any remaining coarse grained HAZ under the buttered layer.

The degree of austenite grain growth can be restricted by using a low heat input. However, precautionary measures may be necessary to avoid the risk of hydrogen-assisted cracking and lack-of-fusion defects. For example, reducing the heat input will almost certainly require a higher preheat temperature to avoid hydrogen-assisted cracking.

The joint design and welding technique adopted should ensure that the weld is free from localised stress concentrations which can arise from the presence of notches. Stress concentrations may be produced in the following situations:

welding with a backing bar a partial penetration weld leaving a root imperfection internal weld imperfections such as lack of sidewall fusion the weld has a poor surface profile, especially sharp weld toesThe weld toes of the capping pass are particularly vulnerable, as the coarse grained HAZ may not have been refined by subsequent passes. In susceptible steel, the last pass should never be deposited on the parent material, but always on the weld metal, so that it will refine the HAZ.

Grinding the weld toes with the preheat maintained has been successfully used to reduce the risk of cracking in 0.5Cr 0.5Mo 0.25V steels.

A general review of geometric shape imperfections - types and causes

Job Knowledge

Part 1. Introduction

In the job knowledge series welding imperfections such as cracks, lack of fusion, penetration and porosity have been discussed. This article looks at those imperfections related to poor geometric shape and will concentrate on the following:

Excess weld metal Undercut

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Overlap Linear misalignment Incompletely filled grooveSuch imperfections might be considered as anomalies in the joint and they will always be present to some degree so that it becomes necessary to separate the acceptable from the unacceptable. This is done by following guidance given by the application standard, which was the basis for the component design, and/or by direction, as set out in the job contract. Examples of standards that might be referred to are:

PD 5500 Specification for unfired fusion welded pressure vessels. BS EN ISO 5817 Welding. Fusion-welded joints in steel, nickel, titanium and their alloys (beam

welding excluded). Quality levels for imperfections AWS D1.1 Structural welding code - SteelExcess weld metal

(also called cap height, overfill or reinforcement)

Fig.1. Excess weld metal

This is weld metal lying outside the plane joining the weld toes. Note that the term 'reinforcement', although used extensively in the ASME/AWS specifications is avoided in Europe as it implies it adds strength to the welded joint, which is rarely the case.

Common causes

This imperfection is formed when excessive weld metal is added to the joint, which is usually a result of poor welder technique for manual processes but may be due to poor parameter selection when the process is mechanised. That is, too much filler metal for the travel speed used. In multi-run welding a poor selection of individual bead sizes can result in a bead build-up pattern that overfills the joint. Different processes and parameters (eg voltage) can result in different excess weld metal shapes.

Acceptance

The acceptability of this imperfection is very dependent on the application in which the product will be used. Most standards have limit, related to material thickness (eg 10%), but also have a maximum upper limits. Both the ratio and the maximum may be related to the severity of service that the component is expected to see. The following table gives examples taken from BS EN ISO 5817.

Excess weld metal limits for quality levels:

Severity of service Moderate, D Stringent, B

Limit (up to maximum) h = 1mm + 0.25 b h = 1mm + 0.1 b

Maximum 10 mm 5 mm

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Transition required smooth smooth

Where: h = height of excess & b = width of bead (see figure 1)

An important reason for limiting the height of excess weld metal is that it represents a non-value added cost. However, it must be remembered that the height of the weld cap influences the resultant toe blend. A sharp transition causes a local stress concentration that can contribute to loss of strength, which is particularly important in fatigue situations. As a result most specifications state that 'smooth transition is required'.

Avoidance

If the imperfection is a result of welder technique then welder retraining is required. For mechanised techniques an increase in travel speed or voltage will help to reduce cap height.

Undercut

Fig.2. Undercut

This is an irregular groove at the toe of a run in the parent metal.

The figure shows undercut at surface of a completed joint but it may also be found at the toes of each pass of a multi-run weld. The latter can result in slag becoming trapped in the undercut region.

Common causes

When arc and gas welding, undercut is probably the most common shape imperfection. With single-sided pipe welds it may also be found at the bore surface. It may also be seen on the vertical face of fillet welds made in the horizontal vertical position.

A wide spreading arc (high arc voltage) with insufficient fill (low current or high travel speed) is the usual cause. However, welder technique, especially when weaving, and the way the welding torch is angled can both cause and be used to overcome undercutting (ie angled to push the weld metal to fill the melted groove). High welding current will also cause undercut - this is generally associated with the need for a high travel speed to avoid overfilling of the joint.

Acceptance

Largely because this imperfection is widespread, most standards permit some level of undercut although they do require that a 'smooth transition is required. The limits in BS EN ISO 5817 range from 0.5mm (stringent) to 1mm (moderate) for thickness (t) greater than 3mm (more stringent limits are required for t 0.5 to 3mm), while AWS D1.1 has a limit of 1mm.

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Measuring undercut can be a problem because of the small size of the imperfection compared with the general environment where there can be mill scale, irregularities in the surface and spatter.

In critical applications the imperfection can be 'corrected' by blend grinding or by depositing an additional weld bead.

Avoidance

This imperfection may be avoided by reducing travel speed and/or the welding current and by maintaining the correct arc length.

Overlap (cold lapping)

Fig.3. Overlap

This is an imperfection at a toe or root of a weld caused by metal flowing on to the surface of the parent metal without fusing to it. It may occur in both fillet and butt welds.

Common causes

This is often caused by poor manipulation of the electrode or welding gun, especially when the weld pool is large and 'cold', where the welder allows gravity to influence the weld shape before solidification. Tightly adherent oxides or scale on the metal surface can also prevent the weld metal fusing with the parent metal to cause the overlap imperfection.

Avoidance

Avoidance is achieved through an acceptable level of welder skill and a reduction in weld pool size (obtained by reducing current or increasing travel speed). Adequate cleaning of the parent plate is also important.

Acceptance

Standards rarely allow the presence of this imperfection, unless the length is short (eg BS EN ISO 5817 for moderate quality level D). Overlap can be very difficult to detect, especially if it is extremely small.

Linear misalignment

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Fig.4 Linear misalignment

(Also known in the USA as high-low).

This imperfection relates to deviations from the correct position/alignment of the joint.

Common causes

This is primarily a result of poor component fit-up before welding, which can be compounded by variations in the shape and thickness of components (eg out of roundness of pipe). Tacks that break during welding may allow the components to move relative to one another, again resulting in misalignment.

Acceptance

The acceptability of this defect is related to the design function of the structure or pipe line either in terms of the ability to take load across the misalignment or because such a step impedes the flow of fluid.

Acceptance varies with the application:

BS EN ISO 5817 relates misalignment to wall thickness but sets maximum limits (eg for material thickness t>3mm and moderate limits of imperfections D, = 0.25 x t, with a maximum of 5mm).

AWS D1.1 allows 10% of the wall thickness up to a maximum of 3mm.

The consequence of linear misalignment can, when welding is carried out from one side, be lack of root or sidewall fusion to give a sharp continuous imperfection along the higher weld face toe. In some situations linear misalignment in the bore of a pipe can lead to in-service problems where turbulence of the carrier fluid in the pipe creates subsequent erosion.

Incomplete filled groove

Incomplete filled groove

This is a continuous, or intermittent, channel in the surface of a weld, running along its length, due to insufficient weld metal.

Common causes

This problem arises when there has been insufficient filler metal (current or wire feed too low or too high a travel speed) so that the joint has not been sufficiently filled. The result is that the thickness of weldment is less than that specified in the design, which could lead to failure.

Acceptance

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Most standards will not accept this type of imperfection, except perhaps over short lengths and even then a smooth transition is required. The designer expects the joint to be adequately filled, but not too much so (see excess weld metal).

Often the presence of this imperfection is an indication of poor workmanship and could suggest that further training is required.

Continuation

Part 2 looks at shape imperfections such as excess penetration and root concavity and highlights shape imperfections related to fillet welded joints.

A general review of the causes and acceptance of shape imperfections - Part 2

Job Knowledge

Click here for Part 1.

This second article on shape imperfections refers mostly to fillet welds but there are two additional butt weld imperfections that require some comment.

Excessive penetration (Excess penetration bead)

Fig.1. Excess penetration

Excess weld metal protruding through the root of a fusion (butt) weld made from one side only.

With pipe welding this type of imperfection may cause effects in the fluid flow that can cause erosion and/or corrosion problems.

Common causes

Penetration becomes excessive when the joint gap is too large, the root faces are too small, the heat input to the joint is too high or a combination of these causes.

Acceptance

The criteria which sets the level of acceptable penetration depends primarily on the application code or specification.

BS 2971 (Class 2 arc welding) requires that the 'penetration bead shall not exceed 3mm for pipes up to and including 150mm bore or 6mm for pipes over 150mm bore'.

BS 2633 (Class 1 arc welding) gives specific limits for smaller diameters pipes, eg for pipe size 25-50mm the maximum allowed bore penetration is 2.5mm.

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ASME B31.3 bases acceptability on the nominal thickness of the weld, for instance, allowing for a thickness range of 13-25mm up to 4mm of protrusion. However, ASME notes that 'more stringent criteria may be specified in the engineering design'.

BS EN ISO 5817 (Quality levels for imperfections), which supersedes BS EN 25817, relates the acceptable protrusion to the width of the under-bead as follows:

Severity of service Moderate, D Stringent, B

Limit (up to maximum) h ≤ 1mm + 1.0 b h ≤ 1mm + 0.2 b

Maximum 5 mm 3 mm

For thicknesses > 3mm where: h = height of excess & b = width of root (see   Fig.1 )

Avoidance

It is important to ensure that joint fit-up is as specified in the welding procedure. If welder technique is the problem then re training is required.

Root concavity (suck-back; underwashing)

Fig.2. Root concavity

A shallow groove that may occur in the root of a butt weld.

Common causes

Root concavity is caused by shrinkage of the weld pool in the through-thickness direction of the weld. Melting of the root pass by the second pass can also produce root concavity.

This imperfection is frequently associated with TIG welding with the most common cause being poor preparation leaving the root gap either too small or, in some cases, too large. Excessively high welding speeds make the formation of root concavity more likely.

Acceptance

The root concavity may be acceptable. This will depend on the relevant standard being worked to. For example:

BS 2971 requires that:a) there is complete root fusionb) the thickness of the weld is not less than the pipe thickness.

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ASME B31.3 requires that the 'total joint thickness, including weld reinforcement, must be greater than the weld thickness'.

BS EN ISO 5817 sets upper limits related to the quality level, eg for thicknesses > 3mm Moderate, (D), h ≤ 0.2t but max 2mm for Stringent, (B), h ≤ 0.05t but max 0.5mm. Furthermore, a smooth transition is required at the weld toes.

In effect the standards require that the minimum design throat thickness of the finished weldment is achieved. If the first two conditions of acceptance are met but the weld face does not have a sufficiently high cap, additional weld metal may be deposited to increase the throat.

Avoidance

It is important to ensure that joint fit-up is as specified in the welding procedure and that the defined parameters are being followed. If welder technique is the problem then retraining is required.

Fillet welded joints

This Section should be read in conjunction with Job Knowledge 66 Fillet welded joints - a review of the practicalities.

Excessive convexity

Fig.3. Excessive convexity

This feature is also covered by the definition for excess weld metal, see Part 1, and may be described as weld metal lying outside the plane joining the weld toes. Note that the term 'reinforcement', although used extensively in the ASME/AWS specifications is avoided in Europe as it implies that excess metal contributes to the strength of the welded joint. This is rarely the case.

Common causes

Poor technique and the deposition of large volumes of 'cold' weld metal.

Acceptance

The idealised design requirement of a 'mitre' fillet weld is often difficult to achieve, particularly with manual welding processes.

BS EN ISO 5817 acceptance is based on a mitre fillet weld shape with a specific design throat and any excess weld metal is measured in relation to this mitre surface. The limits for this imperfection relate the height of the excess metal to the width of the bead with maximum values ranging from 3mm for a stringent

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quality level to 5mm for a moderate quality level. Surprisingly, there is no reference to a 'smooth transition' being required at the weld toes for such weld shape.

AWS D1.1 also has limits relating width to acceptable excess as follows:

Width of weld face Maximum convexity

W ≤ 8mm 2mm

W <8 to W<25mm 3mm

W ≥ 25mm 5mm

Avoidance

Welder technique is the major cause of this problem and training may be required. It is also important to ensure that the parameters specified in the welding procedures specification are adhered to.

Oversize fillet welds (welds with a throat larger than required by the design)

Fig.4. Oversize fillet weld

As discussed in Job Knowledge 66, oversize fillet welds can represent a significant additional cost and loss of productivity.

Common causes

There are some welding related causes, eg high welding current, slow travel speeds, and some supervision related (eg'to be safe make this fillet bigger by x mm').

Acceptance

BS EN ISO 5817 has limits related to the actual throat (eg for stringent quality levels, the actual weld throat [a] may exceed the nominal (design) weld throat [h] by 1+0.15a with a maximum of 3mm. For the moderate quality level (D) the excessive throat thickness is unlimited.

Avoidance

Adhere to the specified welding procedure and parameters and do not add to the specified weld size. Where possible mechanise the welding operation.

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Undersized fillet welds (fillet welds smaller than those specified)

Fig.5. Undersized fillet weld

Common causes

The welding related causes are associated with high welding speeds and low welding currents.

Acceptance

Therefore, it is normally assumed that fillet welds will be at least of the size specified. BS EN SIO 5817 states that limits to insufficient throat thickness are not applicable to processes with proof of greater depth of penetration, therefore a fillet weld with an apparent throat thickness smaller that that prescribed should not be regarded as being imperfect if the actual throat thickness with a compensating greater depth of penetration complies with the nominal value. That is if we can be sure there is good penetration the smaller fillet may be acceptable, however, this should be discussed with the designer of the fabrication. The limits set by the standard.

Relying upon deep penetration to provide the required minimum design throat thickness can be difficult to justify. Penetration is a weld characteristic that is hard to measure directly and reliance must be placed on the stringent control of both the welding process and the welder. Manual welding can rarely be relied upon to provide the required consistency but it is an option with mechanised welding systems.

Imperfection: fillet weldhaving a throat

thickness smaller thanthe nominal value

Quality levels

Moderate D Intermediate C Stringent B

Long imperfections NOT permitted NOT permitted

Short imperfections (see Fig.5) h ≤ 0.3mm+ 0.1 a  

max 2mm max 1mm  

Avoidance

Adhere to the specified welding procedure and parameters. Use sufficient current and appropriate travel speed. Where possible mechanise the welding operation.

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Asymmetric fillet weld (a fillet weld where the legs are of unequal length)

Fig.6. Asymmetric fillet weld

Common causes

Due to incorrect electrode positioning or to gravity pulling the molten pool towards one face of the joint. It is an mainly a problem with fillet welds made in the horizontal/vertical (PB) position.

Acceptance

There are instances where asymmetry may be specified (eg to place the toe stress concentration in a particular region).

BS EN ISO 5817 would, for a 10mm leg length fillet weld (ie 7.1mm throat) allow a difference in leg lengths of about 2.5mm at the stringent quality level and 3.4mm at the moderate quality level. Acceptance is related to the throat thickness.

The consequence of this imperfection is a significant increase in weld volume. Provided the leg length requirement is achieved there would not be a loss of strength. Perhaps this is why, in other standards, a requirement is not specified and the acceptability is left to the inspection personnel to make the 'engineering judgement'!

Poor fit-up

Fig.7. Poor fit-up

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The most common imperfection is an excessive gap between the mating faces of the materials.

Common causes

Poor workshop practice, poor dimensioning and tolerance dimensions on drawings.

Acceptance

A major problem with fillet welds is ensuring the gap between the components is within defined limits. BS EN ISO 5817 specifies the acceptance criteria as follows:

Quality levels

Moderate D Intermediate C Stringent B

h ≤ 1mm + 0.3 a h ≤ 0.5mm + 0.2 a h ≤ 0.5mm + 0.1 a

max 4mm max 3mm max 2mm

Where h = fit-up gap and a = fillet weld design throat

Figure 7 shows that the gap results in a reduction in the leg length on the vertical plate and this, in turn, results in a reduction in the throat thickness of the joint. A 10mm leg length fillet with a root gap of 3mm gives an effective leg of 7mm (a throat of 4.9mm instead of the expected 7mm).

When the application of BS EN ISO 5817 is not required, the guidance of BS EN 1011-2 can be followed, which recommends a maximum gap of 3mm. This standard also states that the size of the fillet weld can be increased to compensate for a large gap.

This discrepancy is addressed within AWS D1.1. which permits a root gap of up to 5mm for material thickness up to 75mm. However, 'if the (joint) separation is greater than 2mm the leg of the fillet weld shall be increased by the amount of the root opening, or the contractor shall demonstrate that the effective throat has been obtained'.

Distortion - Types and causes

Dishing of the steel plate between longitudinal stiffeners can be seen clearly on the bow of this ship (Courtesy MOD)

Job Knowledge

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This article covers several key issues on distortion in arc welded fabrications, especially basic types of and factors affecting the degree of distortion.

What causes distortion?

Because welding involves highly localised heating of joint edges to fuse the material, non-uniform stresses are set up in the component because of expansion and contraction of the heated material. Initially, compressive stresses are created in the surrounding cold parent metal when the weld pool is formed due to the thermal expansion of the hot metal (heat affected zone) adjacent to the weld pool. However, tensile stresses occur on cooling when the contraction of the weld metal and the immediate heat affected zone is resisted by the bulk of the cold parent metal.

The magnitude of thermal stresses induced into the material can be seen by the volume change in the weld area on solidification and subsequent cooling to room temperature. For example, when welding CMn steel, the molten weld metal volume will be reduced by approximately 3% on solidification and the volume of the solidified weld metal/heat affected zone (HAZ) will be reduced by a further 7% as its temperature falls from the melting point of steel to room temperature.

If the stresses generated from thermal expansion/contraction exceed the yield strength of the parent metal, localised plastic deformation of the metal occurs. Plastic deformation causes a permanent reduction in the component dimensions and distorts the structure.

What are the main types of distortion?

Distortion occurs in six main forms:

Longitudinal shrinkage Transverse shrinkage Angular distortion Bowing and dishing Buckling TwistingThe principal features of the more common forms of distortion for butt and fillet welds are shown below:

Contraction of the weld area on cooling results in both transverse and longitudinal shrinkage.

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Non-uniform contraction (through thickness) produces angular distortion in addition to longitudinal and transverse shrinkage.

For example, in a single V butt weld, the first weld run produces longitudinal and transverse shrinkage and rotation. The second run causes the plates to rotate using the first weld deposit as a fulcrum. Hence, balanced welding in a double side V butt joint can be used to produce uniform contraction and prevent angular distortion.

Similarly, in a single side fillet weld, non-uniform contraction produces angular distortion of the upstanding leg. Double side fillet welds can therefore be used to control distortion in the upstanding fillet but because the weld is only deposited on one side of the base plate, angular distortion will now be produced in the plate.

Longitudinal bowing in welded plates happens when the weld centre is not coincident with the neutral axis of the section so that longitudinal shrinkage in the welds bends the section into a curved shape. Clad plate tends to bow in two directions due to longitudinal and transverse shrinkage of the cladding; this produces a dished shape. Dishing is also produced in stiffened plating. Plates usually dish inwards between the stiffeners, because of angular distortion at the stiffener attachment welds (see main photograph).

In plating, long range compressive stresses can cause elastic buckling in thin plates, resulting in dishing, bowing or rippling.

Distortion due to elastic buckling is unstable: if you attempt to flatten a buckled plate, it will probably 'snap' through and dish out in the opposite direction.

Twisting in a box section is caused by shear deformation at the corner joints. This is caused by unequal longitudinal thermal expansion of the abutting edges. Increasing the number of tack welds to prevent shear deformation often reduces the amount of twisting.

How much shall I allow for weld shrinkage?

It is almost impossible to predict accurately the amount of shrinking. Nevertheless, a 'rule of thumb' has been composed based on the size of the weld deposit. When welding steel, the following allowances should be made to cover shrinkage at the assembly stage.

Transverse Shrinkage

Fillet Welds 0.8mm per weld where the leg length does not exceed 3/4 plate thickness

Butt weld 1.5 to 3mm per weld for 60° V joint, depending on number of runs

Longitudinal Shrinkage

Fillet Welds 0.8mm per 3m of weld

Butt Welds 3mm per 3m of weld

Increasing the leg length of fillet welds, in particular, increases shrinkage.

What are the factors affecting distortion?

If a metal is uniformly heated and cooled there would be almost no distortion. However, because the material is locally heated and restrained by the surrounding cold metal, stresses are generated higher than the material yield stress causing permanent distortion. The principal factors affecting the type and degree of distortion, are:

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Parent material properties Amount of restraint Joint design Part fit-up Welding procedureParent material properties

Parent material properties which influence distortion are coefficient of thermal expansion and specific heat per unit volume. As distortion is determined by expansion and contraction of the material, the coefficient of thermal expansion of the material plays a significant role in determining the stresses generated during welding and, hence, the degree of distortion. For example, as stainless steel has a higher coefficient of expansion than plain carbon steel, it is more likely to suffer from distortion.

Restraint

If a component is welded without any external restraint, it distorts to relieve the welding stresses. So, methods of restraint, such as 'strong-backs' in butt welds, can prevent movement and reduce distortion. As restraint produces higher levels of residual stress in the material, there is a greater risk of cracking in weld metal and HAZ especially in crack-sensitive materials.

Joint design

Both butt and fillet joints are prone to distortion. It can be minimised in butt joints by adopting a joint type which balances the thermal stresses through the plate thickness. For example, a double-sided in preference to a single-sided weld. Double-sided fillet welds should eliminate angular distortion of the upstanding member, especially if the two welds are deposited at the same time.

Part fit-up

Fit-up should be uniform to produce predictable and consistent shrinkage. Excessive joint gap can also increase the degree of distortion by increasing the amount of weld metal needed to fill the joint. The joints should be adequately tacked to prevent relative movement between the parts during welding.

Welding procedure

This influences the degree of distortion mainly through its effect on the heat input. As welding procedure is usually selected for reasons of quality and productivity, the welder has limited scope for reducing distortion. As a general rule, weld volume should be kept to a minimum. Also, the welding sequence and technique should aim to balance the thermally induced stresses around the neutral axis of the component.

Distortion - prevention by design

Job knowledge

Strongbacks on girder flange to prevent cross bowing. Courtesy John Allen

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General guidelines are given below as 'best practice' for limiting distortion when considering the design of arc welded structures.

Design principles

At the design stage, welding distortion can often be prevented, or at least restricted, by considering:

elimination of welding weld placement reducing the volume of weld metal reducing the number of runs use of balanced weldingElimination of welding

As distortion and shrinkage are an inevitable result of welding, good design requires that not only the amount of welding is kept to a minimum, but also the smallest amount of weld metal is deposited. Welding can often be eliminated at the design stage by forming the plate or using a standard rolled section, as shown in Fig 1.

Fig. 1 Elimination of welds by:

a) forming the plate; 

b) use of rolled or extruded section 

If possible, the design should use intermittent welds rather than a continuous run, to reduce the amount of welding. For example, in attaching stiffening plates, a substantial reduction in the amount of welding can often be achieved whilst maintaining adequate strength.

Weld placement

Placing and balancing of welds are important in designing for minimum distortion. The closer a weld is positioned to the neutral axis of a fabrication, the lower the leverage effect of the shrinkage forces and the final distortion. Examples of poor and good designs are shown in Fig 2.

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Fig. 2 Distortion may be reduced by placing the welds around the neutral axis

As most welds are deposited away from the neutral axis, distortion can be minimised by designing the fabrication so the shrinkage forces of an individual weld are balanced by placing another weld on the opposite side of the neutral axis. Whenever possible, welding should be carried out alternately on opposite sides, instead of completing one side first. In large structures, if distortion is occurring preferentially on one side, it may be possible to take corrective actions, for example, by increasing welding on the other side to control the overall distortion.

Reducing the volume of weld metal

To minimise distortion, as well as for economic reasons, the volume of weld metal should be limited to the design requirements.

For a single-sided joint, the cross-section of the weld should be kept as small as possible to reduce the level of angular distortion, as illustrated in Fig 3. 

Fig. 3 Reducing the amount of angular distortion and lateral shrinkage by:

a) reducing the volume of weld metal; 

b) using single pass weld

Joint preparation angle and root gap should be minimised providing the weld can be made satisfactorily. To facilitate access, it may be possible to specify a larger root gap and smaller preparation angle. By cutting down the difference in the amount of weld metal at the root and the face of the weld, the degree of angular distortion will be correspondingly reduced. Butt joints made in a single pass using deep penetration have little angular distortion, especially if a closed butt joint can be welded (Fig 3). For example, thin section material can be welded using plasma and laser welding processes and thick section can be welded, in the vertical position, using electrogas and electroslag processes. Although angular distortion can be eliminated, there will still be longitudinal and transverse shrinkage. 

In thick section material, as the cross sectional area of a double-V joint preparation is often only half that of a single-V preparation, the volume of weld metal to be deposited can be substantially reduced. The double-V joint preparation also permits balanced welding about the middle of the joint to eliminate angular distortion.

As weld shrinkage is proportional to the amount of weld metal, both poor joint fit-up and over-welding will increase the amount of distortion. Angular distortion in fillet welds is particularly affected by over-welding. As design strength is based on throat thickness, over-welding to produce a convex weld bead does not increase the allowable design strength but it will increase the shrinkage and distortion.

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Reducing the number of runs

There are conflicting opinions on whether it is better to deposit a given volume of weld metal using a small number of large weld passes or a large number of small passes. Experience shows that for a single-sided butt joint, or a single-side fillet weld, a large single weld deposit gives less angular distortion than if the weld is made with a number of small runs. Generally, in an unrestrained joint, the degree of angular distortion is approximately proportional to the number of passes.

Completing the joint with a small number of large weld deposits results in more longitudinal and transverse shrinkage than a weld completed in a larger number of small passes. In a multi-pass weld, previously deposited weld metal provides restraint, so the angular distortion per pass decreases as the weld is built up. Large deposits also increase the risk of elastic buckling particularly in thin section plate.

Use of balanced welding

Balanced welding is an effective means of controlling angular distortion in a multi-pass butt weld by arranging the welding sequence to ensure that angular distortion is continually being corrected and not allowed to accumulate during welding. Comparative amounts of angular distortion from balanced welding and welding one side of the joint first are shown schematically in Fig 4. The balanced welding technique can also be applied to fillet joints. 

Fig. 4 Balanced welding to reduce the amount of angular distortion

If welding alternately on either side of the joint is not possible, or if one side has to be completed first, an asymmetrical joint preparation may be used with more weld metal being deposited on the second side. The greater contraction resulting from depositing the weld metal on the second side will help counteract the distortion on the first side.

Best practice

The following design principles can control distortion:

eliminate welding by forming the plate and using rolled or extruded sections minimise the amount of weld metal do not over weld use intermittent welding in preference to a continuous weld pass place welds about the neutral axis balance the welding about the middle of the joint by using a double-V joint in preference to a single-

V joint

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Adopting best practice principles can have surprising cost benefits. For example, for a design fillet leg length of 6mm, depositing an 8mm leg length will result in the deposition of 57% additional weld metal. Besides the extra cost of depositing weld metal and the increase risk of distortion, it is costly to remove this extra weld metal later. However, designing for distortion control may incur additional fabrication costs. For example, the use of a double-V joint preparation is an excellent way to reduce weld volume and control distortion, but extra costs may be incurred in production through manipulation of the workpiece for the welder to access the reverse side.

Distortion Control - Prevention by fabrication techniques

Job knowledge

Distortion caused by welding a plate at the centre of a thin plate before welding into a bridge girder section. Courtesy John Allen

Assembly techniques

In general, the welder has little influence on the choice of welding procedure but assembly techniques can often be crucial in minimising distortion. The principal assembly techniques are:

tack welding back-to-back assembly stiffeningTack welding

Tack welds are ideal for setting and maintaining the joint gap but can also be used to resist transverse shrinkage. To be effective, thought should be given to the number of tack welds, their length and the distance between them. With too few, there is the risk of the joint progressively closing up as welding proceeds. In a long seam, using MMA or MIG, the joint edges may even overlap. It should be noted that when using the submerged arc process, the joint might open up if not adequately tacked.

The tack welding sequence is important to maintain a uniform root gap along the length of the joint. Three alternative tack welding sequences are shown in Fig. 1:

a) tack weld straight through to the end of the joint (Fig 1a). It is necessary to clamp the plates or to use wedges to maintain the joint gap during tacking

b) tack weld one end and then use a back stepping technique for tacking the rest of the joint (Fig 1b)

c) tack weld the centre and complete the tack welding by back stepping (Fig 1c).

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Fig. 1. Alternative procedures used for tack welding to prevent transverse shrinkage

a) tack weld straight through to end of joint b) tack weld one end, then use back-step technique for tacking the rest of the joint c) tack weld the centre, then complete the tack welding by the back-step technique 

Directional tacking is a useful technique for controlling the joint gap, for example closing a joint gap which is (or has become) too wide.

When tack welding, it is important that tacks which are to be fused into the main weld are produced to an approved procedure using appropriately qualified welders. The procedure may require preheat and an approved consumable as specified for the main weld. Removal of the tacks also needs careful control to avoid causing defects in the component surface.

Back-to-back assembly

By tack welding or clamping two identical components back-to-back, welding of both components can be balanced around the neutral axis of the combined assembly (Fig. 2a). It is recommended that the assembly is stress relieved before separating the components. If stress relieving is not done, it may be necessary to insert wedges between the components (Fig. 2b) so when the wedges are removed, the parts will move back to the correct shape or alignment.

Fig. 2. Back-to-back assembly to control distortion when welding two identical components

a) assemblies tacked together before welding b) use of wedges for components that distort on separation after welding

Stiffening

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Fig. 3. Longitudinal stiffeners prevent bowing in butt welded thin plate joints

Longitudinal shrinkage in butt welded seams often results in bowing, especially when fabricating thin plate structures. Longitudinal stiffeners in the form of flats or angles, welded along each side of the seam (Fig. 3) are effective in preventing longitudinal bowing. Stiffener location is important: they must be placed at a sufficient distance from the joint so they do not interfere with welding, unless located on the reverse side of a joint welded from one side.

Welding procedure

A suitable welding procedure is usually determined by productivity and quality requirements rather than the need to control distortion. Nevertheless, the welding process, technique and sequence do influence the distortion level.

Welding process

General rules for selecting a welding process to prevent angular distortion are:

deposit the weld metal as quickly as possible use the least number of runs to fill the jointUnfortunately, selecting a suitable welding process based on these rules may increase longitudinal shrinkage resulting in bowing and buckling.

In manual welding, MIG, a high deposition rate process, is preferred to MMA. Weld metal should be deposited using the largest diameter electrode (MMA), or the highest current level (MIG), without causing lack-of-fusion imperfections. As heating is much slower and more diffuse, gas welding normally produces more angular distortion than the arc processes.

Mechanised techniques combining high deposition rates and high welding speeds have the greatest potential for preventing distortion. As the distortion is more consistent, simple techniques such as presetting are more effective in controlling angular distortion.

Welding technique

General rules for preventing distortion are:

keep the weld (fillet) to the minimum specified size use balanced welding about the neutral axis keep the time between runs to a minimum

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Fig. 4. Angular distortion of the joint as determined by the number of runs in the fillet weld

In the absence of restraint, angular distortion in both fillet and butt joints will be a function of the joint geometry, weld size and the number of runs for a given cross section. Angular distortion (measured in degrees) as a function of the number of runs for a 10mm leg length fillet weld is shown in Fig. 4.

If possible, balanced welding around the neutral axis should be done, for example on double sided fillet joints, by two people welding simultaneously. In butt joints, the run order may be crucial in that balanced welding can be used to correct angular distortion as it develops.

Fig. 5. Use of welding direction to control distortion

a) Back-step welding b) Skip welding

Welding sequence

The sequence, or direction, of welding is important and should be towards the free end of the joint. For long welds, the whole of the weld is not completed in one direction. Short runs, for example using the back-step or skip welding technique, are very effective in distortion control (Fig. 5).

Back-step welding involves depositing short adjacent weld lengths in the opposite direction to the general progression (Fig. 5a).

Skip welding is laying short weld lengths in a predetermined, evenly spaced, sequence along the seam (Fig. 5b). Weld lengths and the spaces between them are generally equal to the natural run-out length of one electrode. The direction of deposit for each electrode is the same, but it is not necessary for the welding direction to be opposite to the direction of general progression.

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Best practice

The following fabrication techniques are used to control distortion:

using tack welds to set up and maintain the joint gap identical components welded back to back so welding can be balanced about the neutral axis attachment of longitudinal stiffeners to prevent longitudinal bowing in butt welds of thin plate

structures where there is choice of welding procedure, process and technique should aim to deposit the weld

metal as quickly as possible; MIG in preference to MMA or gas welding and mechanised rather than manual welding

in long runs, the whole weld should not be completed in one direction; back-step or skip welding techniques should be used.

Distortion - corrective techniques

Job knowledge

Local heating of the flange edges to produce curved beams for a bridge structure

Every effort should be made to avoid distortion at the design stage and by using suitable fabrication procedures. As it is not always possible to avoid distortion during fabrication, several well-established corrective techniques can be employed. However, reworking to correct distortion should not be undertaken lightly as it is costly and needs considerable skill to avoid damaging the component.

In this issue, general guidelines are provided on 'best practice' for correcting distortion using mechanical or thermal techniques.

Mechanical techniques

The principal mechanical techniques are hammering and pressing. Hammering may cause surface damage and work hardening.

In cases of bowing or angular distortion, the complete component can often be straightened on a press without the disadvantages of hammering. Packing pieces are inserted between the component and the platens of the press. It is important to impose sufficient deformation to give over-correction so that the normal elastic spring-back will allow the component to assume its correct shape.

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Fig. 1 Use of press to correct bowing in T butt joint

Pressing to correct bowing in a flanged plate is illustrated in Fig. 1. In long components, distortion is removed progressively in a series of incremental pressings; each one acting over a short length. In the case of the flanged plate, the load should act on the flange to prevent local damage to the web at the load points. As incremental point loading will only produce an approximately straight component, it is better to use a former to achieve a straight component or to produce a smooth curvature. 

Best practice for mechanical straightening

The following should be adopted when using pressing techniques to remove distortion:

Use packing pieces which will over correct the distortion so that spring-back will return the component to the correct shape

Check that the component is adequately supported during pressing to prevent buckling Use a former (or rolling) to achieve a straight component or produce a curvature As unsecured packing pieces may fly out from the press, the following safe practice must be

adopted:- bolt the packing pieces to the platen- place a metal plate of adequate thickness to intercept the 'missile'- clear personnel from the hazard area

Thermal techniques

The basic principle behind thermal techniques is to create sufficiently high local stresses so that, on cooling, the component is pulled back into shape. 

Fig. 2 Localised heating to correct distortion

This is achieved by locally heating the material to a temperature where plastic deformation will occur as the hot, low yield strength material tries to expand against the surrounding cold, higher yield strength material. On cooling to room temperature the heated area will attempt to shrink to a smaller size than before heating. The stresses generated thereby will pull the component into the required shape. (See Fig. 2)

Local heating is, therefore, a relatively simple but effective means of correcting welding distortion. Shrinkage level is determined by size, number, location and temperature of the heated zones. Thickness and plate size determines the area of the heated zone. Number and placement of heating zones are largely a question of experience. For new jobs, tests will often be needed to quantify the level of shrinkage.

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Spot, line or wedge-shaped heating techniques can all be used in thermal correction of distortion.

Spot heating

Fig. 3 Spot heating for correcting buckling

Spot heating (Fig. 3), is used to remove buckling, for example when a relatively thin sheet has been welded to a stiff frame. Distortion is corrected by spot heating on the convex side. If the buckling is regular, the spots can be arranged symmetrically, starting at the centre of the buckle and working outwards. 

Line heating

Fig. 4 Line heating to correct angular distortion in a fillet weld

Heating in straight lines is often used to correct angular distortion, for example, in fillet welds (Fig. 4). The component is heated along the line of the welded joint but on the opposite side to the weld so the induced stresses will pull the flange flat.

Wedge-shaped heating

To correct distortion in larger complex fabrications it may be necessary to heat whole areas in addition to employing line heating. The pattern aims at shrinking one part of the fabrication to pull the material back into shape.

Fig. 5 Use of wedge shaped heating to straighten plate

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Apart from spot heating of thin panels, a wedge-shaped heating zone should be used, (Fig. 5) from base to apex and the temperature profile should be uniform through the plate thickness. For thicker section material, it may be necessary to use two torches, one on each side of the plate.

As a general guideline, to straighten a curved plate (Fig. 5) wedge dimensions should be:

1. Length of wedge - two-thirds of the plate width2. Width of wedge (base) - one sixth of its length (base to apex) 

The degree of straightening will typically be 5mm in a 3m length of plate.

Wedge-shaped heating can be used to correct distortion in a variety of situations, (Fig. 6):

1. Standard rolled section which needs correction in two planes (Fig. 6a)2. Buckle at edge of plate as an alternative to rolling (Fig. 6b)

3. Box section fabrication which is distorted out of plane (Fig. 6c)

Fig. 6 Wedge shaped heating to correct distortion a) standard rolled steel section

b) buckled edge of plate

c) box fabrication

General precautions

The dangers of using thermal straightening techniques are the risk of over-shrinking too large an area or causing metallurgical changes by heating to too high a temperature. As a general rule, when correcting distortion in steels the temperature of the area should be restricted to approximately 60° - 650°C - dull red heat.

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If the heating is interrupted, or the heat lost, the operator must allow the metal to cool and then begin again.

Best practice for distortion correction by thermal heating

The following should be adopted when using thermal techniques to remove distortion:

use spot heating to remove buckling in thin sheet structures other than in spot heating of thin panels, use a wedge-shaped heating technique use line heating to correct angular distortion in plate restrict the area of heating to avoid over-shrinking the component limit the temperature to 60° to 650°C (dull red heat) in steels to prevent metallurgical damage in wedge heating, heat from the base to the apex of the wedge, penetrate evenly through the plate

thickness and maintain an even temperatureStandards - application standards, codes of practice and quality levels

Job knowledge

Production at Dennis vehicle manufacturers

Application standards and codes of practice ensure that a structure or component will have an acceptable level of quality and be fit for the intended purpose.

In this document, the requirements for standards on welding procedure and welder approval are explained together with the quality levels for imperfections. It should be noted that the term approval is used in European standards in the context of both testing and documentation. The equivalent term in the ASME standard is qualification.

Application standards and codes

There are essentially three types of standards which can be referenced in fabrication:

Application and design Specification and approval of welding procedures Approval of weldersThere are also specific standards covering material specifications, consumables, welding equipment and health and safety. British Standards are used to specify the requirements, for example, in approving a welding procedure, they are not a legal requirement but may be cited by the Regulatory Authority as a means of satisfying the law. Health and Safety guidance documents and codes of practice may also recommend standards.

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Codes of practice differ from standards in that they are intended to give recommendations and guidance, for example, on the validation of power sources for welding. It is not intended that they should be used as a mandatory, or contractual documents.

Most fabricators will be working to one of the following:

Company or industry specific standards National BS (British Standard) European BS EN (British Standard European Standard) US AWS (American Welding Society) and ASME (American Society of Mechanical Engineers) International ISO (International Standards Organisation)Examples of application codes and standards and related welding procedure and welder/welding operator approval standards are listed in Table 1.

Table 1 Examples of application codes and standards and related welding procedure, welder and welding operator approval standards

    Welding standardApplication Application code/standard Procedure approval Welder approval

Pressure Vessels

PD 5500BS EN 13445 seriesASME B&PV Section III-NB (Nuclear)ASME B&PV Section VIII

BS EN ISO 15614ASME B&PV Section IX

BS EN 287BS EN ISO 9606ASME B&PV Section IX

Process Pipework

BS 2633BS 4677ANSI/ASME B31.1ANSI/ASME B31.3BS 2971

BS EN ISO 15614

ASME B&PV Section I

BS EN ISO 15614-1(if required)

BS EN 287BS EN ISO 9606ASME IXASME IXBS 4872/BS EN 287

Structural Fabrication

AWS D1.1AWS D1.2AWS D1.6BS EN 1011BS 8118

AWS D1.1AWS D1.2AWS D1.6BS EN ISO 15614-1BS EN ISO 15614-2

AWS D1.1AWS D1.2BS EN 287BS EN ISO 9606-2BS 4872

Storage TanksBS EN 14015BS EN 12285API 620/650

BS EN ISO 15614-1, -2BS EN ISO 15614-1, -2ASME IX

BS EN 287BS EN ISO 9606-2ASME IX

Note 1: Reference should be made to the application codes/standards for any additional requirements to those specified in BS EN 287, BS EN ISO 15614, BS EN ISO 9606 and ASME IX.

Note 2: Some BS Standards have not been revised to include the new BS EN standards: BS EN 287/9606 and BS EN ISO 15614 should be substituted, as appropriate, for BS 4871 and BS 4870, respectively, which have been withdrawn.

Note 3: Compliance with the BS EN 13445 series can be used to demonstrate compliance with the Pressure Equipment Directive (PED). Other standards are also acceptable for compliance with the PED, but only provided they take into account the Essential Safety Requirements stated in the PED (ESRs).

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In European countries, national standards are being replaced by EN standards. However, when there is no equivalent EN standard, the National standard can be used. For example, BS EN 287/9606 replace BS 4871 but BS 4872 remains as a valid standard.

Approval of welding procedures and welders

An application standard or code of practice will include requirements or guidelines on material, design of joint, welding process, welding procedure, welder qualification and inspection or may invoke other standards, for example for welding procedure and welder approval tests. The manufacturer will normally be required to approve the welding procedure and welder qualification. The difference between a welding procedure approval and a welder qualification test is as follows:

The welding procedure approval test is carried out by a competent welder and the quality of the weld is assessed using non-destructive and mechanical testing techniques. The intention is to demonstrate that the proposed welding procedure will produce a welded joint which will satisfy the specified requirements of weld quality and mechanical properties.

The welder approval test examines a welder's skill and ability in producing a satisfactory test weld. The test may be performed with or without a qualified welding procedure (note, without an approved welding procedure the welding parameters must be recorded).

The requirements for approvals are determined by the relevant application standard or as a condition of contract (Table 1).

BS EN 287, BS ISO EN 9606 and ASME Section IX would be appropriate for welders on high quality work such as pressure vessels, pressure vessel piping and offshore structures and other products where the consequences of failure, stress levels and complexity mean that a high level of welded joint integrity is essential. In less demanding situations, such as small to medium building frames and general light structural and non- structural work, an approved welding procedure may not be necessary. However, to ensure an adequate level of skill, it is recommended that the welder be approved to a less stringent standard e.g. BS 4872.

'Coded welder' is often used to denote an approved welder but the term is not recognised in any of the standards. However, it is used in the workplace to describe those welders whose skill and technical competence have been approved to the requirements of an appropriate standard.

Quality Acceptance Levels for Welding Procedure and Welder Approval Tests

When welding to application standards and codes, consideration must be given to the imperfection acceptance criteria which must be satisfied. Some standards contain an appropriate section relating to the acceptance levels while others make use of a separate standard. For example, in welding procedure and welder approval tests to BS EN ISO 15614-1 and BS EN ISO 287 Pt1, respectively, reference is made to BS EN ISO 5817. It is important to note that the application standard may specify more stringent imperfection acceptance levels and/or require additional tests to be carried out as part of the welding procedure approval test. For example, for joints which must operate at high temperatures, elevated temperature tensile test may be required whereas for low temperature applications, impact or CTOD tests may be specified.

Guidance on permissible levels of imperfections in arc welded joints in steel are given in BS EN ISO 5817. Production quality, but not fitness-for-purpose, is defined in terms of three levels of quality for imperfections:

Moderate - Level D Intermediate- Level C Stringent - Level B

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The standard applies to most arc welding processes and covers imperfections such as cracks, porosity, inclusions, poor bead geometry, lack of penetration and misalignment.

As the quality levels are related to the types of welded joint and not to a particular component, they can be applied to most applications for procedure and welder approval. The quality levels which are the most appropriate for production joints will be determined by the relevant application standard which may cover design considerations, mode of stressing (e.g. static, dynamic), service conditions (e.g. temperature, environment) and consequences of failure.

When working to the European Standards, the welding procedure, or the welder, will be qualified if the imperfections in the test piece are within the specified limits of Level B except for excess weld metal, excess convexity, excess throat thickness and excess penetration type imperfections when Level C will apply.

Guidance levels for aluminium joints are given in BS EN ISO 10042.

For the American standards ASME Section IX and AWS D1.1, the acceptance levels are contained in the standard. Application codes may specify more stringent imperfection acceptance levels and/or additional tests.

Relevant Standards

American Welding Society, Structural Welding Code, AWS D1.1  ASME Boiler and Pressure Vessel Code, Section IX: Welding Qualifications BS 4872 Approval Testing of Welders when Welding Procedure Approval is not Required BS EN 287-1:2011 Qualification test of welders - fusion welding - Part 1: steels BS EN ISO 9606-2:2004 Qualification test of welders. Fusion welding - Part 2: aluminium and

aluminium alloys. BS EN ISO 15614-1:2004+ A2:2012 Specification and qualification of welding procedures for

metallic materials. Welding procedure test. BS EN ISO 5817:2007 Welding - fusion-welded joints in steel, nickel, titanium and their alloys

(beam welding excluded) - Quality levels for imperfections BS EN ISO 6520-1:2007 Welding and allied processes - Classification of geometric imperfections in

metallic materials BS EN ISO 10042:2005 Welding - Arc welded joints in aluminium and its alloys - Quality levels for

imperfectionsWelding Procedure

Job knowledge

AC TIG welding of aluminium cryogenic pressure vessel Courtesy of Air Products PLC

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For a given application, the main way of ensuring adequate weld quality is to specify the procedure and the skill level of the welding operator. Here, the alternative routes for welding procedure approval are described together with the requirements for welder or welding operator approval.

Routes to welding procedure approval

The key document is the Welding Procedure Specification (WPS) which details the welding variables to be used to ensure a welded joint will achieve the specified levels of weld quality and mechanical properties.

The WPS is supported by a number of documents (eg, a record of how the weld was made, NDE, mechanical test results) which together comprise a welding approval record termed the WPAR (BS EN ISO 15614) or PQR (ASME).

In both the European and ASME standards, there are a number of 'essential variables' specified which, if changed, may affect either weld quality or mechanical properties. Therefore, a change in any of the essentials will invalidate the welding procedure and will require a new approval test to be carried out. The essential variables are detailed in the relevant specification but include material type, welding process, thickness range and sometimes welding position.

Fig. 1. Stages in welding and welder approval

The route followed to produce a WPS in BS EN ISO 15614 and the responsibilities of the manufacturer and the Examiner/Examining Body are shown in Fig.1.

The most common method of gaining approval is to carry out an approval test as described in BS EN ISO 15614-1 (steels) and 15614-2 (aluminium and its alloys). The manufacturer initially drafts a preliminary welding procedure (pWPS) which is used by one of the manufacturer's competent welders to prove that it is capable of producing a welded joint to the specified levels of weld quality and mechanical properties. The welding procedure approval record (WPAR) is a record of this weld. If the WPAR is approved by the Examiner, it is used to finalise one or more WPSs which is the basis for the Work Instructions given to the welder.

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It is noteworthy that the welder carrying out a satisfactory welding procedure approval test is approved for the appropriate range of approval given in the relevant standard (BS EN 287/BS EN ISO 9606, ASME IX or AWS D1.1).

The following options for procedure approval are also possible:

Welding procedure test (BS EN ISO 15614) Approved welding consumable (BS EN ISO 15610) Previous welding experience (BS EN ISO 15611) Standard welding procedure (BS EN ISO 15612) Pre-production welding test (BS EN ISO 15613)The conventional procedure test (as specified in BS EN ISO 15614) does not always need to be carried out to gain approval. But alternative methods have certain limits of application regarding, for example, welding processes, materials and consumables as specified in the appropriate application standard or contract agreement.

The welding procedure test method of approval is often a mandatory requirement of the Application Standard. If not, the contracting parties can agree to use one of the alternative methods. For example, a welding procedure specification can be approved in accordance with the requirements of BS EN ISO 15611 (previous experience) on condition that the manufacturer can prove, with appropriate documentation, that the type of joint has previously been welded satisfactorily.

The American standard, ASME IX requires a welding procedure test (PQR) but AWS D1.1 will allow the use of pre-qualified procedures within the limits detailed in the specification.

Welder approval

The welder approval test is carried out to demonstrate that the welder has the necessary skill to produce a satisfactory weld under the conditions used in production as detailed in the approved WPS or Work Instruction. As a general rule, the test piece approves the welder not only for the conditions used in the test but also for all joints which are considered easier to weld.

As the welder's approval test is carried out on a test piece which is representative of the joint to be welded, it is independent of the type of construction. The precise conditions, called 'essential variables', must be specified in the approval test, eg material type, welding process, joint type, dimensions and welding position. The extent of approval is not necessarily restricted to the conditions used for the test but covers a group of similar materials or a range of situations which are considered easier to weld.

It is important to note that a number of Amendments and Corrigenda have now been issued which affect the range of approval (see list of Relevant Standards).

In BS EN 287/BS EN ISO 9606, the certificate of approval testing is issued under the sole responsibility of the Examiner/Examining Body. The welder approval certificate remains valid subject to the requirements of the application standard. In BS EN 287/BS EN ISO 9606, it can be extended at six monthly intervals by the employer for up to two years provided the welder has been successfully welding similar joints. After two years, prolongation of the welder's qualification will need approval of the Examiner who will require proof that his or her performance has been of the required standard during the period of validity. As the Examiner will normally examine the company's records on the welder's work and tests as proof that he has maintained his skill, it is essential that work records are maintained by the company.

It should also be noted that BS EN 287/BS EN ISO 9606 requires records of tests, ie half yearly documentation about X-ray or ultrasonic inspections or test reports on fracture tests must be maintained

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with the welder's approval certificate (tests on production welds will satisfy this requirement). Failure to comply will necessitate a retest.

American standards have similar requirements although the extent of approval of the welding variables are different to those of BS EN 287/BS EN ISO 9606.

Welding operator approval

When required by the contract or application standard, the welding operators responsible for setting up and/or adjustment of fully mechanised and automatic equipment must be approved but the personnel operating the equipment do not need approval. In clarifying the term 'welding operator', personnel who are using the equipment (loading and unloading robotic equipment or operating a resistance welding machine) do not require approval.

As specified in BS EN 1418, approval of operators of equipment for fusion welding and resistance weld equipment setters can be based on:

welding a procedure test pre-production welding test or production test production sample testing or a function test.It should be noted that the methods must be supplemented by a functional test appropriate to the welding unit. However, a test of knowledge relating to welding technology which is the equivalent of 'Job knowledge for welders' in BS EN 287/BS EN ISO 9606 is recommended but not mandatory.

Prolongation of the welding operator approval is generally in accordance with the requirements of BS EN 287/BS EN ISO 9606. The welding operator's approval remains valid for two years providing the employer/welding co-ordinator confirms that there has been a reasonable continuity of welding work (period of interruption no longer than six months) and there is no reason to question the welding operator's knowledge.

The validity of approval may be prolonged for further periods of two years by the examiner / examining body providing there is proof of production welds of the required quality, and appropriate test records maintained with the operator's certificate.

When working to ASME IX, operators for both mechanised and automatic welding equipment require approval. The essential variables are different to those in welder approval.

Relevant Standards

BS EN 287-1:2004 Qualification test of welders - fusion welding - Part 1: steels

BS EN ISO 15614Specification and qualification of welding procedures for metallic materials - Welding Procedure test. Part 1 Arc and gas welding of steels and arc welding of nickel and nickel alloys.

Part 2: Arc welding of aluminium and its alloys Part 3: Fusion and pressure welding of non-alloyed and low-alloyed cast irons Part 4: Finishing welding of aluminium castings Part 5: Arc welding of titanium, zirconium and their alloys Part 6: Arc and gas welding of copper and its alloys Part 7: Overlay welding Part 8: Welding of tubes to tube-plate joints Part 9: Underwater hyperbaric wet welding

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Part 10: Hyperbaric dry welding Part 11: Electron and laser beam welding Part 12: Spot, seam and projection welding Part 13: Resistance butt and flash welding BS EN ISO 9606-2: 2004Qualification test of welders - fusion welding. Part 2: Aluminium and aluminium alloys

BS EN ISO 15612: 2004Specification and qualification of welding procedures for metallic materials - Qualification by adoption of a standard welding procedure.

BS EN ISO 15610: 2003Specification and qualification of welding procedures for metallic materials - Qualification based on tested welding consumables.

BS EN ISO 15611: 2003Specification and qualification of welding procedures for metallic materials - Qualification based on previous welding experience.

BS EN 1418 : 1998 Welding personnel - Approval testing of welding operators for fusion welding and resistance weld setters for fully mechanised and automatic welding of metallic materials

A review of the application of weld symbols on drawings - Part 1

Job Knowledge

Weld symbols have been used for many years and are a simple way of communicating design office details to a number of different industrial shop floor personnel such as welders, supervisors, and inspectors. Subcontractors are often required to interpret weld symbols on engineering drawings, from perhaps the main contractor or client. It is essential that everyone should have a full understanding of weld symbol requirements to ensure that the initial design requirement is met.

There are a number of standards which relate to weld symbols including British, European, International and American (American Welding Society) standards. Most of the details are often similar or indeed, the same, but it is essential that everyone concerned knows the standard to be used. One of the first requirements therefore is:

Which standard?

The UK has traditionally used BS 499 Part 2. This standard has now been superseded by BS EN 22553, however in many welding and fabrication organisations there will be old drawings used that make reference to out of date standards such as BS 499 Pt 2.

BS EN 22553 is almost identical to the original ISO 2553 standard on which it was based. Therefore we can say, for at least this article's scope, there are no significant differences, but it is essential that the reader consults the specific standard. The American system is also similar in many respects but will not be covered here.

Basic requirements

All the standards have the same requirements in relation to the following items:

Arrow line and arrow head Reference line

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The arrow line can be at any angle (except 180 degrees) and can point up or down. The arrow head must touch the surfaces of the components to be joined and the location of the weld. Any intended edge preparation or weldment is not shown as an actual cross sectional representation, but is replaced by a line. The arrow also points to the component to be prepared with single prepared components. See Figs. 1-4.

Fig. 1.

Fig. 3.

Fig. 2.

Fig. 4.

Symbol types

To the basic set-up of the arrow and reference line, the design draughtsperson can apply the appropriate symbol, or symbols for more complex situations.

The symbols, in particular for arc and gas welding, are often shown as cross sectional representations of either a joint design or a completed weld. Simple, single edge preparations are shown in Fig. 5.

For resistance welding, a spot weld and seam weld are shown in Fig. 6:

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Fig. 5.

Fig. 6.

Joint and/or weld shape

The above examples can be interpreted as either the joint details alone or the completed weld, however, for a finished weld it is normal to find that an appropriate weld shape is specified. Using the examples above, there are a number of options and methods to specify an appropriate weld shape or finish.

Butt welded configurations would normally be shown as a convex profile (Fig.7 'a', 'd' and 'f') or as a dressed-off weld as shown in 'b' and 'c'. Fillet weld symbols are always shown as a 'mitre' fillet weld (a right angled triangle) and a convex or concave profile can be superimposed over the original symbol's mitre shape. See Fig. 7.

Fig. 7.

Part 2 of this explanation of weld symbols covering more complex situations will appear in the next issue.

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A review of the application of weld symbols on drawings - Part 2

Job Knowledge

Part 1 of this article which appeared in the May/June issue of Connect, dealt with the most basic weld symbols as they appear on engineering drawings. As previously mentioned, it is essential that all concerned in any project are aware of which Standard is being applied.

Weld sizing

In order that the correct size of weld can be applied, it is common to find numbers to either the left or to the right of the symbol.

For fillet welds, numbers to the left of the symbol indicate the design throat thickness, leg length, or both design throat thickness and leg length requirements. Figure 1 gives examples of symbols used in different Standards.

Fig.1

For fillet welds:

Superseded BS499 Pt 2 gives

a = design throat thicknessb = leg length

ISO 2553/EN 22553 requirements

a = design throat thicknessz = leg lengths = penetration throat thickness

For butt joints and welds, an S with a number to the left of a symbol refers to the depth of penetration as shown inFig.2.

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Fig.2

When there are no specific dimensional requirements specified for butt welds on a drawing using weld symbols, it would normally be assumed that the requirement is for a full penetration butt weld ( Fig.3).

Fig.3

Numbers to the right of a symbol or symbols relate to the longitudinal dimension of welds, eg for fillets, the number of welds, weld length and weld spacing for non-continuous welds, as Fig.4.

Fig.4

On fillet welded joints made from both sides, a staggered weld can be shown by placing a 'Z' through the reference line ( Fig.5).

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Fig.5

Supplementary symbols

Weld symbols indicate the type of preparation to use or the weld type. However, there may still be occasions where other information is required. The basic information can therefore be added to in order to provide further details as shown in Figs.6, 7 and 8.

Fig.6

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Fig.7

Fig.8

Weld all round

For a Rectangular Hollow Section (RHS) welded to a plate, for example:

Weld in the field or on site

The box attached to the arrow can be used to contain, or point to, other information.

Welding process type

ISO 4063 gives welding processes specific reference numbers. As shown in Fig.9 the appropriate process number is placed in the tail of the arrow. Other processes are given a unique number. In this example, 135 refers to MAG welding.

Fig.9

There are a number of additional symbols given in the Standards ( eg ISO 22553) which refer to additional welding or joint requirements. Figure 10 shows the requirement for a sealing run.

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Fig.10

Compound joints/welds

A compound weld could be a 'T' butt weld which requires fillet welds to be added to increase the throat thickness as shown in Fig.11.

Fig.11

The broken reference line

Fig.12

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The main feature that distinguishes weld symbol standards is that for ISO 2553 and BS EN 22553, there is an additional feature of a broken reference line.

This method is used when a weldment or weld preparation needs to be specified on the 'other side' of the arrow as shown in Fig.12.

Any symbol that is used to show a joint or weld type feature on the other side of the arrow line is always placed on a dotted line.

BS 499 and AWS require symbols to be placed above the reference line (which indicate the other side) or below the reference line (indicating the arrow side).

Summary

Weld symbols are a very useful way of communicating welding requirements from the design office to the shop floor.

It is essential that the 'rules' of the standard used are correctly applied by drawing office personnel. However, it is also important that shop floor personnel are able to read and understand the details of weld symbols.

Much of this requirement can be met by reference to the standard being used within the organisation and by the drawing office personnel considering the needs of the end user such as the welders, welding supervisors, welding inspection personnel and welding engineers in order to minimise costly mistakes due to misinterpretation.

Training of all personnel in the correct use of weld symbol specifications also plays an important role in ensuring that weld symbols are both correctly applied and correctly read.

Fillet welded joints - a review of the practicalities

Job Knowledge

Fillet welded joints such as tee, lap and corner joints are the most common connection in welded fabrication. In total they probably account for around 80% of all joints made by arc welding.

It is likely that a high percentage of other joining techniques also use some form of a fillet welded joint including non-fusion processes such as brazing, braze welding and soldering. The latter techniques are outside the scope of this article.

Although the fillet weld is so common, there are a number of aspects to be considered before producing such a weld. This article will review a number of topics that relate to fillet welded joints and it is hoped that even the most seasoned fabricator or welding person will gain from this article in some way.

Common joint designs for fillet welds are shown below in Fig.1.

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Fig 1. Common joint designs for fillet welds

Fillet weld features

ISO 2553 (EN 22553) uses the following notation as Figs.2 and 3 show.

a = throat thicknessz = leg lengths = deep penetration throat thicknessl = length of intermittent fillet

Fig. 2. Mitre fillet

Fig 3. Deep penetration fillet

Fillet weld shapes

Over specified fillet welds or oversized fillet welds

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Fig 4. Weld sizes in relation to the required leg lengths or throat thickness

One of the greatest problems associated with fillet welded joints is achieving the correct weld size in relation to the required leg lengths or throat thickness (Fig.4).

The designer may calculate the size and allow a 'safety factor' so that the weld specified on the fabrication drawing is larger than is required by design considerations.

The weld size is communicated by using an appropriate weld symbol.

In the UK the weld size is frequently specified by referring to the leg length 'z' in ISO 2553 where the number gives the weld size in millimetres as shown in Fig.5.

Fig 5. Weld size specification (UK)

In Europe, it is more common to find the design throat thickness, 'a' specified ( Fig.6).

Fig 6. Weld size specification (Europe)

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Once the drawing has been issued to the shop floor, it is usual to find an additional safety factor also being applied on by the welder or inspector. It is also common to hear 'add a bit more it will make it stronger'.

The outcome is an oversized weld with perhaps an 8mm leg length rather than the 6mm specified by the designer. This extra 2mm constitutes an increase in weld volume of over 80%.

This coupled with the already over specified weld size from the designer's 'safety factor' may lead to a weld that is twice the volume of a correctly sized fillet weld.

By keeping the weld to the size specified by the drawing office, faster welding speeds can be achieved, therefore increasing productivity, reducing overall product weight, consumable consumption and consumable cost.

The other benefit is that, in the case of most arc welding processes, a slight increase in travel speed would in most cases see an increase in root penetration so that the actual throat thickness is increased:

An oversized weld is therefore very costly to produce, may not have 'better strength' and is wasteful of welding consumables and may see other fabrication problems including excessive distortion.

Lap joints welded with fillet welds.

As discussed earlier, oversized welds are commonplace and the lap joint is no exception. The designer may specify a leg length that is equal to the material thickness as in Fig.7.

Fig 7. Lap joint - leg length specification

Strength considerations may mean that the fillet weld size need not be anywhere near the plate thickness. In practice the weld may also be deficient in other ways for example:

Fig 8. Example showing an undersized fillet weld

Due to melting away of the corner of the upper plate (Fig.8), the vertical leg length is reduced meaning that the design throat has also been reduced; therefore an undersized weld has been created. Care is therefore needed to ensure that the corner of the upper plate is not melted away. Ideally the weld should be some 0.5-1mm clear of the top corner (Fig.9).

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Fig 9. Ideally the weld should be 0.5-1mm clear of the top corner

It may be the designer may therefore specify a slightly smaller leg length compared to the thickness of the component.

To compensate for this reduction in throat thickness it may be necessary to specify a deep penetration fillet weld. This amount of additional penetration would need to be confirmed by suitable weld tests. Additional controls may also be needed during production welding to ensure that this additional penetration is being achieved consistently.

In addition to the reduction in throat thickness there is the potential for additional problems such as overlap at the weld toe due to the larger weld pool size (Fig.10) or an excessively convex weld face and consequential sharp notches at the weld toe (Fig.11). 

Fig 10. Overlap at the weld toe due to the larger weld pool size

Fig 11. Excessively convex weldface and consequential sharp notches at the weld toe

Both the potential problems shown in Figs.10 and 11 could adversely influence the fatigue life of the welded joint due to the increased toe angle, which acts as a greater stress concentration.

Poor fit-up can also reduce the throat thickness as in Fig.12. The corner of the vertical component has been bevelled in the sketch in an exaggerated manner to illustrate the point.

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Fig 12. Throat thickness may be reduced by poor fit-up

Summary

Fillet welded joints are not only the most frequently used weld joints but are also one of the most difficult to weld with any real degree of consistency. Fillet welds require a higher heat input than a butt joint of the same thickness and, with less skilled welders this can lead to lack of penetration and/or fusion defects that cannot be detected by visual examination and other NDT techniques.

Fillet welded joints are not always open to NDT or are indeed time consuming to many non-destructively testing techniques such as radiography or ultrasonic testing and the results are often difficult to interpret. Inspection methods such as visual inspection, magnetic particle inspection and penetrant inspection are surface examination techniques only and with visual inspection, much of the effort is expended in measuring the size of the weld rather than identifying other quality aspects.

Fillet welded joints are therefore much more difficult to weld and inspect. Often the welds that are produced are larger than they need to be or they may be of a poor shape which can adversely influence their service performance.

To overcome these difficulties, designers need to specify accurately the most appropriate throat size and welding personnel should strive to achieve the specified design size. Welders also need to be adequately trained and sufficiently skilled to be capable of maintaining an acceptable weld quality.

Mechanical testing - Tensile testing, Part 1

Job Knowledge

Mechanical testing is carried out to produce data that may be used for design purposes or as part of a material joining procedure or operator acceptance scheme. The most important function may be that of providing design data since it is essential that the limiting values that a structure can withstand without failure are known.

Fig.1. Typical tensile testing machine

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Inadequate control of the material properties by the supplier, or incompetent joining procedures and operatives are, however, equally crucial to the supply of a product that is safe in use. An example of this dual role of mechanical testing is the tensile test that may be used either to determine the yield strength of a steel for use in design calculations or to ensure that the steel complies with a material specification's strength requirements.

Mechanical tests may also be divided into quantitative or qualitative tests. A quantitative test is one that provides data that will be used for design purposes, a qualitative test where the results will be used for making comparisons - hardness or Charpy-V tests - for example as a 'go/no go test' such as the bend test.

Mechanical property data are obtained from a relatively small number of standard tests and these will be covered over the next several articles. These will include tensile and toughness tests, the tests used for welding procedure and welder approval and those used for the determination of in-service properties.

Tensile testing

As mentioned earlier the tensile test is used to provide information that will be used in design calculations or to demonstrate that a material complies with the requirements of the appropriate specification - it may therefore be either a quantitative or a qualitative test.

The test is made by gripping the ends of a suitably prepared standardised test piece in a tensile test machine and then applying a continually increasing uni-axial load until such time as failure occurs. Test pieces are standardised in order that results are reproducible and comparable as shown in Fig 2.

Fig.2. Standard shape tensile specimens

Specimens are said to be proportional when the gauge length, L0, is related to the original cross sectional area, A0, expressed as L0 =k√A0 . The constant k is 5.65 in EN specifications and 5 in the ASME codes. These give gauge lengths of approximately 5x specimen diameter and 4x specimen diameter respectively - whilst this difference may not be technically significant it is important when claiming compliance with specifications.

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Fig.3. Stress/strain curve

Both the load (stress) and the test piece extension (strain) are measured and from this data an engineering stress/strain curve is constructed, Fig.3. From this curve we can determine:

a) the tensile strength, also known as the ultimate tensile strength, the load at failure divided by the original cross sectional area where the ultimate tensile strength (U.T.S.), σmax = Pmax /A0 , where Pmax = maximum load, A0 = original cross sectional area. In EN specifications this parameter is also identified as 'Rm';

b) the yield point (YP), the stress at which deformation changes from elastic to plastic behaviour ie below the yield point unloading the specimen means that it returns to its original length, above the yield point permanent plastic deformation has occurred, YP or σy = Pyp /A0 where Pyp = load at the yield point. In EN specifications this parameter is also identified as 'Re ';

c) By reassembling the broken specimen we can also measure the percentage elongation, El% how much the test piece had stretched at failure where El% = (Lf - L0 /Lo ) x100 where Lf = gauge length at fracture and L0 = original gauge length. In EN specifications this parameter is also identified as 'A' ( Fig.4a).

d) the percentage reduction of area, how much the specimen has necked or reduced in diameter at the point of failure where R of A% =(A0 - Af /A0 ) x 100 where Af = cross sectional area at site of the fracture. In EN specifications this parameter is also identified as 'Z', ( Fig.4b).

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Fig.4: a) Calculation of percentage elongation, b) Calculation of percentage reduction of area

(a) and (b) are measures of the strength of the material, (c) and (d) indicate the ductility or ability of the material to deform without fracture.

The slope of the elastic portion of the curve, essentially a straight line, will give Young's Modulus of Elasticity, a measure of how much a structure will elastically deform when loaded.

A low modulus means that a structure will be flexible, a high modulus a structure that will be stiff and inflexible.

To produce the most accurate stress/strain curve an extensometer should be attached to the specimen to measure the elongation of the gauge length. A less accurate method is to measure the movement of the cross-head of the tensile machine.

The stress strain curve in Fig.3 shows a material that has a well pronounced yield point but only annealed carbon steel exhibits this sort of behaviour. Metals that are strengthened by alloying, by heat treatment or by cold working do not have a pronounced yield and some other method must be found to determine the 'yield point'.

This is done by measuring the proof stress ( offset yield strength in American terminology), the stress required to produce a small specified amount of plastic deformation in the test piece.

The proof stress is measured by drawing a line parallel to the elastic portion of the stress/strain curve at a specified strain, this strain being a percentage of the original gauge length, hence 0.2% proof, 1% proof (see Fig.5).

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Fig.5. Determination of proof (offset yield) strength

For example, 0.2% proof strength would be measured using 0.2mm of permanent deformation in a specimen with a gauge length of 100mm. Proof strength is therefore not a fixed material characteristic, such as the yield point, but will depend upon how much plastic deformation is specified. It is essential therefore when considering proof strengths that the percentage figure is always quoted. Most steel specifications use 0.2% deformation, RP0.2 in the EN specifications.

Some materials such as annealed copper, grey iron and plastics do not have a straight line elastic portion on the stress/strain curve. In this case the usual practice, analogous to the method of determining proof strength, is to define the 'yield strength' as the stress to produce a specified amount of permanent deformation.

Part 2 of this series on mechanical testing will cover welding procedure approval tensile testing.

Mechanical testing - Tensile testing Part II

Welding procedure approval for tensile testing.

Job Knowledge

200

To approve a butt welding procedure most specifications such as ISO 15614 and ASME IX require tensile tests to be carried out.

These are generally cross joint (CJ) tensile tests of square or rectangular cross section that, as the name suggests, are oriented across the weld so that both parent metals, both heat affected zones (HAZs) and the weld metal itself are tested ( Fig.1). The excess weld metal in the cap of the weld may be left in-situ or machined off.

Fig.1. Square or rectangular cross joint tensile test piece

While it is possible to measure the yield strength, the elongation and the reduction of area of CJ specimens the fact that there are at least three different areas with dissimilar mechanical properties makes such measurements inaccurate and unreliable, although this is sometimes carried out purely for information purposes.

The specifications mentioned above require the UTS and the position of the fracture only to be recorded. The cross joint strength is usually required to exceed the minimum specified UTS of the parent metal. In most situations the weld metal is stronger than the parent metal - it is overmatched - so that failure occurs in the parent metal or the HAZ at a stress above the specified minimum.

In cases where the weld and/or the HAZs are weaker than the parent metal - welded age-hardened or cold worked aluminium alloys are a good example - this is covered in most specifications.

The designer must also take this into account in design calculations and provide some method of compensating for this loss of strength.

The tensile testing of flat plate butt welds presents few problems of specimen shape but those machined from a pipe butt joint are not flat and this curvature can affect the results. In the context of welding procedure approval testing, this is not significant since the test is used only for the determination of the UTS and the position of the fracture. For more accurate results the test piece may be waisted and may be machined flat as illustrated in Fig.2.

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Fig.2. Flat cross joint tensile specimen machined from tube

It may be necessary to machine a number of specimens through the thickness of a weld, particularly on very thick joints where the capacity of the tensile machine is insufficient to pull a full thickness specimen, Fig.3.

Fig.3. Multiple cross joint specimens machined from thick plate

To test a small diameter tube, a solid bar is inserted in the bore of the tube to prevent the tube collapsing when the sample is clamped into the tensile machine.

Most weld testing is carried out with CJ specimens but longitudinally oriented specimens are useful particularly where the weld metal or the HAZ is very strong but ductility is low.

In a CJ specimen the parent metal can yield and finally fail without the weld metal or the HAZ experiencing any significant amount of deformation whereas in a longitudinal test piece the load is shared more equally.

A brittle weld or HAZ will not elongate with the parent metal but will crack, with the cracks opening, but not necessarily propagating into the parent metal, as testing proceeds.

The testing described above is that required by the welding procedure approval specifications. These provide no assurance that the welds in a structure will be suitable for their purpose such as elevated or

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cryogenic service and many application standards such as PD 5500 Unfired Pressure Vessels, and ASME VIII Pressure Vessels, require additional tests.

Since the strength of a metal falls as the temperature rises these specifications require elevated temperature tensile tests to be carried out at the maximum design temperature.

These tests are required to be carried out on the weld metal only and use a longitudinally orientated round cross section specimen from which an accurate measurement of the proof strength can be obtained.

Many application standards such as PD 5500 require tests additional to those required by, for example, ISO 15614-1. This must be remembered when procedure approval documentation is submitted for approval by the inspecting authority or the client.

Validity of tensile data

The samples taken are assumed to be representative of the bulk of the material but this is not always the case.

Tensile strength of a casting, for instance, is often determined from a specimen machined from a riser and this will have a grain size different from that of the bulk of the casting.

A rolled steel plate will be found to have different properties in the longitudinal, transverse and through thickness directions. Material specifications such as BS EN 10028, Flat Products in Steel for Pressure Purposes, therefore, require the tensile test to be taken transverse to the rolling direction so that the steel is tested across the 'grain' - the lower strength, lower ductility direction.

The size of a product can also influence the properties as, during heat treatment, the section thickness will affect the cooling rate with slower cooling rates, and hence softer structures, at the centre of thicker sections. This is dealt with in material standards by specifying what is known as the 'limiting ruling section', the maximum diameter of bar at which the required mechanical properties can be achieved at the centre.

In addition to variations of the properties due to the shape of the specimens and the testing temperature, the rate of loading will also affect the results.

Figure 4 shows how the tensile strength increases but ductility decreases as the testing speed is increased. The speed of the cross head of the tensile machine therefore needs to be controlled and BS EN 10002 specifies a stress rate range of 6MPa per second to 60MPa per second. The ASTM specifications have similar - but not identical - requirements.

Fig.4. Effect of speed of testing on strength and ductility

Needless to say, calibration of testing equipment to guarantee operation within acceptable parameters is mandatory.

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Relevant specifications

BS EN 10002 Methods of tensile testing of metallic materials.BS EN 876 Destructive tests on welds in metallic materials - longitudinal tensile test.BS EN 895 Destructive tests on welds in metallic materials - transverse tensile test.BS EN ISO 7500-1 Tension/compression testing machines. verification and calibration of the force measuring system.ASTM A370 Mechanical testing of steel products.ASTM E8 Tension testing of metallic materials.ASTM B557 Tension testing wrought and cast aluminium and magnesium alloy products.

Mechanical testing - notched bar or impact testing

Job Knowledge

Before looking at impact testing let us first define what is meant by 'toughness' since the impact test is only one method by which this material property is measured.

Toughness is, broadly, a measure of the amount of energy required to cause an item - a test piece or a bridge or a pressure vessel - to fracture and fail. The more energy that is required then the tougher the material.

The area beneath a stress/strain curve produced from a tensile test is a measure of the toughness of the test piece under slow loading conditions. However, in the context of an impact test we are looking at notch toughness, a measure of the metal's resistance to brittle or fast fracture in the presence of a flaw or notch and fast loading conditions.

It was during World War II that attention was focused on this property of 'notch toughness' due to the brittle fracture of all-welded Liberty ships, then being built in the USA. From this work the science of fracture toughness developed and gave rise to a range of tests used to characterise 'notch toughness' of which the Charpy-V test described in this article is one.

There are two main forms of impact test, the Izod and the Charpy test.

Both involve striking a standard specimen with a controlled weight pendulum travelling at a set speed. The amount of energy absorbed in fracturing the test piece is measured and this gives an indication of the notch toughness of the test material.

These tests show that metals can be classified as being either 'brittle' or 'ductile'. A brittle metal will absorb a small amount of energy when impact tested, a tough ductile metal a large amount of energy.

It should be emphasised that these tests are qualitative, the results can only be compared with each other or with a requirement in a specification - they cannot be used to calculate the fracture toughness of a weld or parent metal, such as would be needed to perform a fitness for service assessment. Fracture toughness tests that can be used in this way are covered in other Job Knowledge articles (Job Knowledge 76 & 77). The Izod test is rarely used these days for weld testing having been replaced by the Charpy test and will not be discussed further in this article.

The Charpy specimen may be used with one of three different types of notch, a 'keyhole', a 'U' and a 'V'. The keyhole and U-notch are used for the testing of brittle materials such as cast iron and for the testing of plastics. The V-notch specimen is the specimen of choice for weld testing and is the one discussed here.

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The current British Standard for Charpy testing is BS EN ISO 148-1:2009 and the American Standard is ASTM E23. The standards differ only in the details of the strikers used. The standard Charpy-V specimen, illustrated in Fig.1. is 55mm long, 10mm square and has a 2mm deep notch with a tip radius of 0.25mm machined on one face.

Fig.1. Standard Charpy-V notch specimen

To carry out the test the standard specimen is supported at its two ends on an anvil and struck on the opposite face to the notch by a pendulum as shown in Fig.2. The specimen is fractured and the pendulum swings through, the height of the swing being a measure of the amount of energy absorbed in fracturing the specimen. Conventionally three specimens are tested at any one temperature, see Fig.3, and the results averaged.

Fig.2. Charpy testing machine

Fig.3. Schematic Charpy-V energy and % age crystallinity curves

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A characteristic of carbon and low alloy steels is that they exhibit a change in fracture behaviour as the temperature falls with the failure mode changing from ductile to brittle.

If impact testing is carried out over a range of temperatures the results of energy absorbed versus temperature can be plotted to give the 'S' curve illustrated in Fig.3.

This shows that the fracture of these types of steels changes from being ductile on the upper shelf to brittle on thelower shelf as the temperature falls, passing through a transition region where the fracture will be mixed.

Many specifications talk of a transition temperature, a temperature at which the fracture behaviour changes from ductile to brittle. This temperature is often determined by selecting, quite arbitrarily, the temperature at which the metal achieves an impact value of 27 Joules - see, for example the impact test requirements of EN 10028 Part 2 Steel for Pressure Purposes.

What the curve shows is that a ductile fracture absorbs a greater amount of energy than a brittle fracture in the same material. Knowing the temperature at which the fracture behaviour changes is therefore of crucial importance when the service temperature of a structure is considered - ideally in service a structure should operate at upper shelf temperatures.

The shape of the S curve and the positions of the upper and lower shelves are all affected by composition, heat treatment condition, whether or not the steel has been welded, welding heat input, welding consumable and a number of additional factors. All the factors must be controlled if good notch toughness is required. This means that close control of the welding parameters is essential if impact testing is a specification requirement. Austenitic stainless steels, nickel and aluminium alloys do not show this change in fracture behaviour, the fracture remaining ductile even to very low temperatures. This is one reason why these types of alloys are used in cryogenic applications.

In addition to the impact energy there are two further features that can be measured and may be found as a requirement in some specifications. These are percentage crystallinity and lateral expansion.

The appearance of a fracture surface gives information about the type of fracture that has occurred - a brittle fracture is bright and crystalline, a ductile fracture is dull and fibrous.

Percentage crystallinity is therefore a measure of the amount of brittle fracture, determined by making a judgement of the amount of crystalline or brittle fracture on the surface of the broken specimen.

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Lateral expansion is a measure of the ductility of the specimen. When a ductile metal is broken the test piece deforms before breaking, a pair of 'ears' being squeezed out on the side of the compression face of the specimen, as illustrated in Fig 4. The amount by which the specimen deforms is measured and expressed as millimetres of lateral expansion. ASME B31.3 for example requires a lateral expansion of 0.38mm for bolting materials and steels with a UTS exceeding 656N/mm2, rather than specifying an impact value.

 

Fig.4 Lateral expansion

The next article in this series will look at the testing of welds, how the impact strength can be affected by composition and microstructure and some of its limitations and disadvantages.

Notched bar or impact testing. Part II

Job Knowledge

Mechanical testing - Notched bar or impact testing - Part I

The previous article looked at the method of Charpy-V impact testing and the results that can be determined from carrying out a test. This next part looks at the impact testing of welds and some of the factors that affect the transition temperature such as composition and microstructure. Within such a short article, however, it will only be possible to talk in the most general of terms.

Welding can have a profound effect on the properties of the parent metal and there may be many options on process selection, welding parameters and consumable choice that will affect impact strength.

Many application standards therefore require impact testing to be carried out on the parent metal, the weld metal and in the heat affected zone as illustrated in Fig.1 which is taken from BS PD 5500 Annex D. The standards generally specify a minimum impact energy to be achieved at the minimum design temperature and to identify from where the specimens are to be taken. This is done in order to quantify the impact energy of the different microstructures in the weld metal and the HAZs to ensure that, as far as possible, the equipment will be operating at upper shelf temperatures where brittle fracture is not a risk.

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Fig.1. PD5500 App D. location of Charpy specimens in weld HAZ

These application standards may be supplemented by client specifications that impose additional and more stringent testing requirements, as shown in Fig.2 taken from an oil industry specification for offshore structures.

Fig.2. Offshore client requirements

The positioning of the specimens within a weld is extremely important both in terms of the specimen location and the notch orientation. A specimen positioned across the width of a multi-pass arc weld will probably include more than one weld pass and its associated HAZs. Quite a small movement in the position of the notch can therefore have a significant effect on the impact values recorded during a test. Positioning a notch precisely down the centre line of a single pass of a submerged arc weld can give extremely low impact values!

Testing the heat affected zone also has problems of notch position since in a carbon or low alloy steel there will be a range of microstructures from the fusion line to the unaffected parent metal. Many welds also use a 'V' preparation as illustrated above and this, coupled with the narrow HAZ, means that a single notch may sample all of these structures. If the impact properties of specific areas in the HAZ need to be determined then a 'K' or single bevel preparation may be used.

The standard specimen is 10mm x 10mm square - when a weld joint is thicker than 10mm the machining of a standard size specimen is possible. When the thickness is less than this and impact testing is required it becomes necessary to use sub-size specimens.

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Many specifications permit the use of 10mm x 7.5mm, 5mm and 2.5mm thickness (notch length) specimens. There is not a simple relationship between a 10mm x 10mm specimen and the sub-size specimens - a 10mm x 5mm specimen does not have half the notch toughness of the full size test piece. As the thickness decreases the transition temperature also decreases, as does the upper shelf value, illustrated in Fig.3 and this is recognised in the application standards.

Fig.3. Effect of size on transition temperature and upper shelf values

In a carbon or low alloy steel the lowest impact values are generally to be found close to the fusion line where grain growth has taken place.

Coarse grains generally have low notch toughness, one reason why heat input needs to be controlled to low levels if high notch toughness is required.

For example, EN ISO 15614 Pt. 1 requires Charpy-V specimens to be taken from the high heat input area of a procedure qualification test piece and places limits on any increase in heat input. Certain steels may also have an area some distance from the fusion line that may be embrittled so some specifications require impact tests at a distance of 5mm from the fusion line.

Charpy-V tests carried out on rolled products show that there is a difference in impact values if the specimens are taken parallel or transverse to the rolling direction. Specimens taken parallel to the rolling direction test the metal across the 'grain' of the steel and have higher notch toughness than the transverse specimens - one reason why pressure vessel plates are rolled into cylinders with the rolling direction oriented in the hoop direction.

In a carbon or low alloy steel the element that causes the largest change in notch toughness is carbon with the transition temperature being raised by around 14°C for every 0.1% increase in carbon content.

An example of how this can affect properties is the root pass of a single sided weld. This often has lower notch toughness than the bulk of the weld as it has a larger amount of parent metal melted into it - most parent metals have higher carbon content than the filler metal and the root pass therefore has a higher carbon content than the bulk of the weld.

Sulphur and phosphorus are two other elements that both reduce notch toughness, one reason why steel producers have been working hard to reduce these elements to as low a level as possible. It is not uncommon for a good quality modern steel to have a sulphur content less than 0.005%.

Of the beneficial elements, manganese and nickel are possibly the two most significant, the nickel alloy steels forming a family of cryogenic steels with the 9% nickel steel being capable of use at temperatures

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down to -196°C. aluminium is also beneficial at around 0.02% where it has the optimum effect in providing a fine grain size.

Lastly, let us have a brief look at some of the other factors that can affect the impact values. These are concerned with the quality of the specimen and how the test is conducted.

It goes without saying that the specimens must be accurately machined, the shape of the tip of the notch being the most important feature. A blunted milling cutter or broach will give a rounded notch tip and this in turn will give a false, high impact value. Checking the tip radius on a shadowgraph is one simple way of ensuring the correct tip shape. Correct positioning of the specimen on the anvil is most important and this can be done using a specially designed former.

The last point concerns the testing of specimens at temperatures other than at room temperature. When testing at sub-zero temperatures the length of time taken to remove the specimen from the cooling bath, position it on the anvil and test it is most important. EN875 requires this to be done within five seconds otherwise the test piece temperature will rise making the test invalid - referring back to the impact energy vs temperature curve in the previous article will show why.

Relevant Specifications

BS 131 Part 4 Calibration of Impact Testing Machines for metals.BS 131 Part 5 Determination of CrystallinityBS 131 Part 6 Method for Precision Determination of Charpy-V Impact EnergyBS 131 Part 7 Specification for Verification of Precision Test MachinesEN 875   Destructive Tests on Welds in Metallic Materials - Impact TestsEN 10045 Part 1 Test MethodEN 10045 Part 2 Verification of Impact Testing MachinesASTM E23-O2A Standard Test Methods for Notched Bar Impact Testing of Metallic Materials.

Bend testing

Job Knowledge

The bend test is a simple and inexpensive qualitative test that can be used to evaluate both the ductility and soundness of a material. It is often used as a quality control test for butt-welded joints, having the advantage of simplicity of both test piece and equipment.

No expensive test equipment is needed, test specimens are easily prepared and the test can, if required, be carried out on the shop floor as a quality control test to ensure consistency in production.

The bend test uses a coupon that is bent in three point bending to a specified angle.

The outside of the bend is extensively plastically deformed so that any defects in, or embrittlement of, the material will be revealed by the premature failure of the coupon.

The bend test may be free formed or guided.

The guided bend test is where the coupon is wrapped around a former of a specified diameter and is the type of test specified in the welding procedure and welder qualification specifications. For example, it may be a requirement in ASME IX, ISO 9606 and ISO 15614 Part 1.

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As the guided bend test is the only form of bend test specified in welding qualification specifications it is the only one that will be dealt with in this article.

Typical bend test jigs are illustrated in Fig.1(a) and 1(b).

Fig.1(a) shows a guided bend test jig that uses a male and a female former, the commonest form of equipment

Fig.1(b) shows a wrap-around guided bend test machine that works on the same principles as a plumber's pipe bender

The strain applied to the specimen depends on the diameter of the former around which the coupon is bent and this is related to the thickness of the coupon 't', normally expressed as a multiple of 't' eg 3t, 4t etc.

The former diameter is specified in the test standard and varies with the strength and ductility of the material - the bend former diameter for a low ductility material such as a fully hard aluminium alloy may be as large as 8t. An annealed low carbon steel on the other hand may require a former diameter of only 3t. The angle of bend may be 90°, 120° or 180° depending on the specification requirements.

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Fig.2 Material over 12mm thick is normally tested using the side bend test that tests the full section thickness

On completion of the test the coupon is examined for defects that may have opened up on the tension face. Most specifications regard a defect over 3mm in length as being cause for rejection.

For butt weld procedure and welder qualification testing the bend coupons may be oriented transverse or parallel to the welding direction.

Below approximately 12mm material thickness transverse specimens are usually tested with the root or face of the weld in tension. Material over 12mm thick is normally tested using the side bend test that tests the full section thickness, Fig.2.

Where the material thickness is too great to permit the full section to be bent the specifications allow a number of narrower specimens to be taken provided that the full material thickness is tested. Conventionally, most welding specifications require two root and two face bend coupons or four side bends to be taken from each butt welded test piece.

The transverse face bend specimen will reveal any defects on the face such as excessive undercut or lack of sidewall fusion close to the cap. The transverse root bend is also excellent at revealing lack of root fusion or penetration. The transverse side bend tests the full weld thickness and is particularly good at revealing lack of side-wall fusion and lack of root fusion in double-V butt joints. This specimen orientation is also useful for testing weld cladding where any brittle regions close to the fusion line are readily revealed.

Longitudinal bend specimens are machined to include the full weld width, both HAZs and a portion of each parent metal. They may be bent with the face, root or side in tension and are used where there is a difference in mechanical strength between the two parent metals or the parent metal and the weld. The test will readily reveal any transverse defects but it is less good at revealing longitudinally oriented defects such as lack of fusion or penetration.

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Whilst the bend test is simple and straightforward to perform there are some features that may result in the test being invalid.

In cutting the coupon from the test weld the effects of the cutting must not be allowed to affect the result. Thus it is necessary to remove any HAZ from flame cutting or work hardened metal if the sample is sheared.

It is normal to machine or grind flat the face and root of a weld bend test coupon to reduce the stress raising effect that these would have. Sharp corners can cause premature failure and should be rounded off to a maximum radius of 3mm.

The edges of transverse bend coupons from small diameter tubes will experience very high tensile stresses when the ID is in tension and this can result in tearing at the specimen edges.

Weld joints with non-uniform properties such as dissimilar metal joints or where the weld and parent metal strengths are substantially different can result in 'peaking' of the bend coupon. This is when most of the deformation takes place in the weaker of the two materials which therefore experiences excessive localised deformation that may result in premature failure.

A dissimilar metal joint where one of the parent metals is very high strength is a good example of where this may occur and similar peaking can be seen in fully hard welded aluminium alloy joints.

In these instances the roller bend test illustrated in Fig.1(b) is the best method of performing a bend test as each component of the coupon is strained by a similar amount and peaking is to a great extent eliminated.

Related Specifications

BS EN 910 Destructive Tests on Welds in Metallic Materials - Bend TestsASME IX Welding and Brazing QualificationsASTM E190-92 Guided bend Test for Ductility of Welds

Hardness Testing Part 1

Job Knowledge

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The hardness of a material can have a number of meanings depending upon the context, which in the case of metals generally means the resistance to indentation. There are a number of test methods of which only the Brinell, Vickers and portable hardness testing will be covered in this article.

Brinell Hardness Test

The Brinell test was devised by a Swedish researcher at the beginning of the 20th century. The test comprises forcing a hardened steel ball indentor into the surface of the sample using a standard load as shown in Fig.1(a). The diameter/load ratio is selected to provide an impression of an acceptable diameter. The ball may be 10, 5 or 1mm in diameter, the load may be 3000, 750 or 30kgf, The load, P, is related to the diameter, D by the relationship P/D2 and this ratio has been standardised for different metals in order that test results are accurate and reproducible. For steel the ratio is 30:1 - for example a 10mm ball can be used with a 3000kgf load or a 1mm ball with a 30kgf load. For aluminium alloys the ratio is 5:1. The load is applied for a fixed length of time, usually 30 seconds. When the indentor is retracted two diameters of the impression, d1 and d2 , are measured using a microscope with a calibrated graticule and then averaged as shown in Fig.1(b).

Fig.1. Brinell Hardness Test

The Brinell hardness number (BHN) is found by dividing the load by the surface area of the impression. There is a somewhat tedious calculation that can be carried out to determine the hardness number but it is more usual and far simpler to refer to a set of standard tables from which the Brinell hardness number can be read directly.

The Brinell test is generally used for bulk metal hardness measurements - the impression is larger than that of the Vickers test and this is useful as it averages out any local heterogeneity and is affected less by surface roughness. However, because of the large ball diameter the test cannot be used to determine the hardness variations in a welded joint for which the Vickers test is preferred. Very hard metals, over 450BHN may also cause the ball to deform resulting in an inaccurate reading. To overcome this limitation a tungsten carbide ball is used instead of the hardened steel ball but there is also a hardness limit of 600BHN with this indentor.

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Vickers Hardness Test

The Vickers hardness test operates on similar principles to the Brinell test, the major difference being the use of a square based pyramidal diamond indentor rather than a hardened steel ball. Also, unlike the Brinell test, the depth of the impression does not affect the accuracy of the reading so the P/D2 ratio is not important. The diamond does not deform at high loads so the results on very hard materials are more reliable. The load may range from 1 to 120kgf and is applied for between 10 and 15 seconds.

The basic principles of operation of the Vickers hardness test are illustrated in Fig.2 where it can be seen that the load is applied to the indentor by a simple weighted lever. In older machines an an oil filled dash pot is used as a timing mechanism - on more modern equipment this is done electronically.

Fig.2. Schematic principles of operation of Vickers hardness machine

As illustrated in Fig.3(b) two diagonals, d1 and d2 , are measured, averaged and the surface area calculated then divided into the load applied. As with the Brinell test the diagonal measurement is converted to a hardness figure by referring to a set of tables. The hardness may be reported as Vickers Hardness number (VHN), Diamond Pyramid Number (DPN) or, most commonly, Hvxx where 'xx' represents the load used during the test.

Fig.3. Vickers hardness test

As mentioned earlier, the Vickers indentation is smaller than the Brinell impression and thus far smaller areas can be tested, making it possible to carry out a survey across a welded joint, including individual runs and the heat affected zones. The small impression also means that the surface must be flat and perpendicular to the indentor and should have a better than 300 grit finish.

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Errors in Hardness Testing

There are many factors that can affect the accuracy of the hardness test. Some of these such as flatness and surface finish have already been mentioned above but it is worth re-emphasising the point that flatness is most important - a maximum angle of approximately ± 1° would be regarded as acceptable.

To achieve the required flatness tolerance and surface finish surface grinding or machining may be necessary. The correct load must be applied and to achieve this there must be no friction in the loading system otherwise the impression will be smaller than expected - regular maintenance and calibration of the machine is therefore essential. The condition of the indentor is crucial - whilst the Vickers diamond is unlikely to deteriorate with use unless it is damaged or loosened in its mounting by clumsy handling, the Brinell ball will deform over a period of time and inaccurate readings will result. This deterioration will be accelerated if a large proportion of the work is on hard materials. The length of time that the load is applied is important and must be controlled.

The specimen dimensions are important - if the test piece is too thin the hardness of the specimen table will affect the result. As a rule of thumb the specimen thickness should be ten times the depth of the impression for the Brinell test and twice that of the Vickers diagonal. Similarly, if the impression is too close to the specimen edge then low hardness values will be recorded - again as a rule the impression should be some 4 to 5 times the impression diameter from any free edge. Performing hardness testing on cylindrical surfaces eg pipes and tubes, the radius of curvature will affect the indentation shape and can lead to errors. It may be necessary to apply a correction factor - this is covered in an ISO specification, ISO 6507 Part 1.

The specimen table should be rigidly supported and must be in good condition - burrs or raised edges beneath the sample will give low readings. Impact loading must be avoided. It is very easy to force the indentor into the specimen surface when raising the table into position. This can strain the equipment and damage the indentor. Operator training is crucial and regular validation or calibration is essential if hardness rest results are to be accurate and reproducible.

Hardness Testing Part 2

Job Knowledge

Part 1

216

The previous article dealt with the conventional Vickers and Brinell hardness tests. This second article reviews micro-hardness and portable hardness testing. The investigation of metallurgical problems in welds often requires the determination of hardness within a very small area or on components in service or too large to be able to test in a laboratory environment.

Micro-hardness testing may be carried out using any one of three common methods and, as with the macro-hardness tests, measure the size of the impression produced by forcing an indentor into the specimen surface under a dead load, although many of the new test machines use a load cell system.

The three most common tests are the Knoop test, the Vickers test and the ultrasonic micro-hardness test.

The Knoop test uses a pyramidal indentor that gives an elongated diamond shaped impression with an aspect ratio of around 7:1, the Vickers test uses the pyramidal indentor described in the previous article (January/February 2005).

The Knoop test is rarely used in Europe where the Vickers test is the preferred method. The loads used for the tests vary from 1gmf to 1kgf and produce impressions that need to be measured by using a microscope with magnifications of up to 100X, although modern machines may be equipped with an image analysis system that enables the process to be automated.

The ultrasonic hardness test does not rely upon measuring the size of an impression. Instead, the test uses a Vickers diamond attached to the end of a metal rod. The rod is vibrated at its natural frequency by a piezoelectric converter and then brought into contact with the specimen surface under a small load. The resonant frequency is changed by the size of the impression produced and this change can be measured and converted to a hardness value.

The size of the impression is extremely small and the test may be regarded as non-destructive since it is non-damaging in most applications.

The micro-hardness test has a number of applications varying from being a metallurgical research tool to a method of quality control. The test may be used to determine the hardness of different micro-constituents in a metal, as shown in Fig.1. Where an impression would be damaging, for instance on a finished product, micro-hardness tests, particularly the ultrasonic test, may be used for quality control purposes. Micro-hardness testing also finds application in the testing of thin foils, case hardened items and decarburised components.

Fig.1. Micro-hardness test

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Portable hardness tests may be used where the component is too large to be taken to the testing machine or in on-site applications. It is useful on-site, for example, for checking that the correct heat treatment has been carried out on welded items or that welded joints comply with the hardness limits specified by NACE for sour service. There are three principal methods - dynamic rebound, Brinell or Vickers indentation or ultrasonic testing.

The Leeb hardness test uses dynamic rebound where a hammer is propelled into the test piece surface and the height of the rebound is measured. This gives a measure of the elasticity of the material and hence its hardness.

This type of test is typified by the 'Equotip' test, Fig.2, a trademark of Proceq SA. The Equotip tester comprises a hand-held tube that contains a spring loaded hammer. The device is cocked by compressing the hammer against the spring, the device is then positioned vertically on the test surface and the release button is pressed. The hammer strikes the surface, rebounds and the result displayed digitally. Generally the average of five readings is taken.

Fig.2. Equotip test

To obtain a valid result, the position of the device, the flatness of the surface and the flexibility of the component all affect the accuracy of the results. Needless to say the skill and experience of the operator is one of the key factors in producing accurate hardness figures. The results are generally converted to give a hardness in Vickers or Brinell units.

The other type of portable hardness test in common use is the ultrasonic method described above. Commercially available machines are typified by the Microdur unit supplied by GE Inspection Technologies as shown in Fig.3. This type of equipment is electronically based and can be programmed to give hardness readings of any type - Vickers, Brinell, or Rockwell. Needless to say, any of these methods of hardness testing require regular calibration of the equipment, fully trained operators and well prepared surfaces.

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Fig.3. Ultrasonic testing using a Microdur unit

Although there are several different methods of hardness testing the results can be compared and converted. The ASTM specification E140 contains conversion tables for metals - ferritic and austenitic steels, nickel alloys, copper and brass- for converting Vickers to Brinell or Rockwell or vice versa.

To end this article on hardness testing let us look at the significance of the results.

Hardness is related to tensile strength - multiplying the Vickers hardness number of a carbon steel by 3.3 will give the approximate ultimate tensile strength in N/mm2 . A hardness traverse across a weld and its HAZs will therefore reveal how the tensile strength varies. In carbon or low alloy steels a hardness of above approximately 380HV suggests that the hard brittle microstructure, martensite, has been formed leading to the possibility of cold cracking during fabrication or brittle fracture in service. This fact has been recognised in the specification EN ISO 15614 Part 1 so that a maximum hardness of 380HV is permitted on a hardness traverse of a macro-section from a carbon steel procedure qualification test.

Relevant Specifications

ASTM E 10 Brinell Hardness of Metallic MaterialsASTM E 140 Hardness Conversion Tables for Metals.ASTM E 110 Portable Hardness Testing.ASTM E 384 Microhardness Testing of Metallic Materials.ASTM E 103 Rapid Indentation Hardness Testing.ASTM E 18 Rockwell Hardness Testing.ASTM E 92 Vickers Hardness of Metallic Materials.

Fatigue testing

Job Knowledge

Fatigue as a specific failure mechanism has been recognised since the early part of the nineteenth century but it was the development of rail travel that resulted in a major increase of interest in this type of fracture.

The premature failure of wagon axles led to Wohler in Germany investigating fatigue failure under rotating loading. This led to the design of the first standardised test - a reversing stress rotating specimen, illustrated in Fig.1.

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Fig.1. Wohler rotating fatigue test

There are many mechanisms that can lead to failure but fatigue is perhaps one of the most insidious since it can lead to a catastrophic failure with little or no warning - one well known example being the failure of the Comet aircraft in the 1950s.

Failure can occur at a fluctuating load well below the yield point of the metal and below the allowable static design stress. The number of cycles at which failure occurs may vary from a couple of hundreds to millions. There will be little or no deformation at failure and the fracture has a characteristic surface, as shown in Fig.2.

Fig.2. Typical fatigue crack fracture surface

The surface is smooth and shows concentric rings, known as beach marks, that radiate from the origin; these beach marks becoming coarser as the crack propagation rate increases. Viewing the surface on a scanning electron microscope at high magnification shows each cycle of stress causes a single ripple. The component finally fails by a ductile or brittle overload.

Fatigue cracks generally start at changes in section or notches where the stress is raised locally and, as a general rule, the sharper the notch the shorter the fatigue life - one reason why cracks are so damaging.

There are two stages in the process of fatigue cracking - a period of time during which a fatigue crack is nucleated and a second stage where the crack grows incrementally leaving the ripples described above. In an unwelded component the bulk of the life is spent in initiating a fatigue crack with a shorter period spent in crack propagation.

An unwelded ferritic steel component exhibits an endurance limit - a stress below which fatigue cracking will not initiate and failure will therefore not occur. This is not the case with most non-ferrous metals or with welded joints - these have no clearly defined endurance limit.

The reason for this is that in arc welded joints there is an 'intrusion' - a small defect at the toe of the weld, perhaps only some 0.1mm deep. Provided that the applied stress is sufficiently large a crack will begin to propagate within an extremely short period of time. The endurance limit for a welded joint is therefore

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dependent on the intrusion size that does not result in crack propagation at the applied stress range. In the case of a welded joint, therefore, a fatigue limit - a 'safe life' is specified, often the stress to cause failure at 2x106 or 107 cycles.

During fatigue the stress may alternate about zero, may vary from zero to a maximum or may vary about some value above - or below - zero.

To quantify the effect of these varying stresses fatigue testing is carried out by applying a particular stress range and this is continued until the test piece fails. The number of cycles to failure is recorded and the test then repeated at a variety of different stress ranges.

This enables an S/N curve, a graph of the applied stress range, S, against N, the number of cycles to failure, to be plotted as illustrated in Fig.3. This graph shows the results of testing a plain specimen and a welded component. The endurance limit of the plain specimen is shown as the horizontal line - if the stress is below this line the test piece will last for an infinite number of cycles. The curve for the welded sample, however, continues to trend down to a point where the stress range is insufficient to cause a crack to propagate from the intrusion.

Fig.3. S/N curves for welded and unwelded specimens

By testing a series of identical specimens it is possible to develop S/N curves. In service however, there will be variations in stress range and frequency. The direction of the load may vary, the environment and the shape of the component will all affect the fatigue life, as explained later in this article.

When designing a test to determine service performance it is therefore necessary to simulate as closely as possible these conditions if an accurate life is to be determined. In order to enable the fatigue life to be calculated when the stress range varies in this random manner, the Palmgren-Miners cumulative damage rule is used.

This rule states that, if the life at a given stress is N and the number of cycles that the component has experienced is a smaller number, n, then the fatigue life that has been used up is n/N.

If the number of cycles at the various stress ranges are then added together - n1/N1 + n2/N2 + n3/N3 + n4/N4 etc - the fatigue life is used up when the sum is of all these ratios is 1. Although this does not give a precise estimate of fatigue life, Miners rule was generally regarded as being safe. This method, however, has now been superseded with the far more accurate approach detailed in the British Standard BS 7608.

The design of a welded joint has a dominant effect on fatigue life. It is therefore necessary to ensure that a structure that will experience fatigue loading in the individual joints has adequate strength. The commonest method for determining fatigue life is to refer to S/N curves that have been produced for the relevant weld designs.

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The design rules for this range of joint designs were first developed by TWI and incorporated with the bridge code BS 5400 in 1980 and then into the industry design rules for offshore structures. Further refinements and improvements finally resulted in the publication of BS 7608 Code of practice for fatigue design and assessment of steel structures. This standard will be looked at in more detail in a future article.

Fatigue testing - Part 2

Job Knowledge

Part 1 Part 3

The article in the September/October issue of Connect established some basic facts about fatigue and the statement was made that a welded joint exhibited no clearly established fatigue limit as in an unwelded component. In this article we will be looking at some of the reasons for this behaviour.

It should be mentioned that, in service, few structures experience purely static loads and that most will be subjected to some fluctuations in applied stresses and may therefore be regarded as being fatigue loaded. Motorway gantries, for example, are buffeted by the slipstream from large lorries and offshore oil rigs by wave action. Process pressure vessels will experience pressure fluctuations and may also be thermally cycled.

If these loads are not accounted for in the design, fatigue failure may occur in as few as a couple of tens of cycles or several million and the result may be catastrophic when it does.

Fatigue failures can occur in both welded and unwelded components, the failure usually initiating at any changes in cross section - a machined groove, a ring machined onto a bar or at a weld. The sharper the notch the greater will be its effect on fatigue life.

The effect of a change in section is illustrated in Fig.1, where it can be seen that the stress is locally raised at the weld toe. The illustration shows a bead-on-plate run but a full penetration weld will show the same behaviour.

Fig.1. Stress concentrating effect of a change in thickness

In addition, misalignment and/or distortion of the joint will cause the applied stress to be further increased, perhaps by introducing bending in the component, further reducing the expected fatigue life. A poorly shaped weld cap with a sharp transition between the weld and the parent metal will also have an adverse effect on fatigue performance.

In addition to these geometrical features affecting fatigue life there is also the small intrusion at the weld toe, mentioned in the last article and illustrated in Fig.2. In an unwelded component the bulk of the fatigue life is spent in initiating the fatigue crack with a smaller proportion spent in the crack propagating through the structure. In a welded component the bulk of the fatigue life is spent in propagating a crack. The consequences of this difference in behaviour are illustrated in Fig.3.

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Fig.2. Weld toe intrusion

Fig.3. Effect of stress concentration on fatigue life

This shows that this small intrusion reduces the fatigue life of a fillet welded joint by a factor of perhaps 10 compared with that of an unwelded item and some eight times that of a sample with a machined hole. The other consequence is that fatigue cracks in welded joints almost always initiate at the toe of a weld, either face or root.

It may be thought that the use of a higher strength material will be of benefit in increasing fatigue life. The rate of crack propagation, however, is determined by Young's Modulus - a measure of the elastic behaviour of the metal - and not simply by tensile strength.

Alloying or heat treatment to increase the strength of a metal has very little effect on Young's Modulus and therefore very little effect on crack propagation rates. Since the bulk of a welded component's life is spent in propagating a crack, strength has little or no influence on the fatigue life of a welded item. There is thus no benefit to be gained by using high strength alloys if the design is fatigue limited. This is illustrated in Fig.4 which shows the benefits of increasing the ultimate tensile strength of a steel if the component is unwelded or only machined but how little effect this has on the life of a welded item.

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Fig.4. Effect of increase in tensile strength on fatigue life

One additional feature in welded joints that set them apart from unwelded or machined items is the presence of residual tensile stress.

In a welded component there will be stresses introduced into the structure by, for example, assembly stress. These stresses are long range reaction stresses and from a fatigue point of view have little effect on fatigue life.

Of far greater significance with respect to fatigue are the short range stresses introduced into the structure by the expansion and contraction of material close to and within the welded joint. Whilst the actual level of residual stress will be affected by such factors as tensile strength, joint type and size and by run size and sequence, the peak residual stress may be regarded as being of yield point magnitude. The implications of this are that it is the stress range that determines fatigue life and not the magnitude of the nominal applied stress.

Even if the applied stress range is wholly compressive and there is apparently no fluctuating tensile stress to cause a crack to form and grow, the effect of welding residual stress is to make the structure susceptible to fatigue failure. This is illustrated in Fig.5, where it can be seen that, irrespective of the applied stress, the effective stress range is up to the level of residual stress at the welded joint.

Fig.5. Effect of residual stress on stress range

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It would seem reasonable, therefore, that a post-weld stress relief treatment would be of benefit to the fatigue life by reducing the residual stresses to low levels. This is only true, however, where the applied stress range is partly or wholly compressive. If the applied stress range is all tensile, research has shown that as-welded and stress relieved components have almost identical fatigue performances with only a marginal improvement in the stress relieved joints.

This is the result of the bulk of the fatigue life of a welded joint being spent in crack propagation where propagation rates are only marginally affected by mean stress. It may be difficult therefore to justify the cost of stress relief if the only criterion is that of improving fatigue life.

The methods of determining fatigue performance of welded joints, as detailed in BS 7608, and how fatigue performance can be improved will be dealt with in the next Connect article.

Fatigue testing Part 3

Job Knowledge

Part 1   Part 2

What will have become obvious from the previous two articles on fatigue is that a welded joint behaves in a radically different way from an unwelded item, even if this item contains a significant stress raiser.

The last article, number 79, made the statement that a welded joint exhibits no clearly defined fatigue limit, the limit varying dependent upon the joint type and weld quality. It is vitally important to understand this if fatigue analysis of welded joints is to be carried out.

As mentioned earlier, rules for the design of components subject to fatigue loading were produced by TWI and these were incorporated into the design rules in BS 5400, the British bridge design code. These rules were later adopted by the offshore industry for offshore structures and adaptations of these rules now appear in many other specifications such as BS PD 5500 Unfired pressure vessels and BS 8118 Structural use of aluminium.

The basis of all the rules is a system whereby various joint designs are assigned a 'classification' related to the joint's fatigue performance. Fig.1 is an example of how this classification has been formalised in BS 7608 - the same or similar methods will be found in other application standards.

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Fig.1. Examples of joint classification from BS 7608

In BS 7608 each joint type is assigned a classification letter. For example, a plate butt weld with cap and root ground flush is class 'C', an undressed plate butt weld class 'D' and a fillet weld class 'F' ( Fig.2).

Fig.2. Effects of joint classification on fatigue life

For each classification a fatigue curve has been developed and from these curves the design life can be predicted. This is obviously an over-simplification of what can be a very complicated task -the forces acting on a joint arising from changes in temperature, changes in internal or external pressure, vibration, externally applied fluctuating loads etc can be complex and difficult to determine.

Whilst the joint design has a major effect on design life and is the basis for calculating service performance, the weld quality also has a decisive effect - any fatigue analysis assumes that the welds are of an acceptable quality and comply with the inspection acceptance standards. However, in practice it is not always possible

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to guarantee a 'perfect' weld and cracks, lack of fusion, slag entrapment and other planar defects may be present, reducing the fatigue life, perhaps catastrophically.

Other less obvious features will also have an adverse effect. Excessive cap height or a poorly shaped weld bead will raise the stress locally and reduce the design life; misalignment may cause local bending with a similar effect. Good welding practices, adherence to approved procedures and competent and experienced staff will all help in mitigating these problems.

In some applications an as-welded joint will not have a sufficient design life and some method of improving the fatigue performance needs to be found. There are a number of options available. The first and perhaps simplest is to move the weld from the area of highest stress range, the next is to thicken up the component or increase the weld size. Note that, as mentioned in the earlier article, using a higher strength alloy will not improve the fatigue life.

Local spot heating to induce compressive stresses at the weld toes will also help, although this needs very accurate positioning of the heated area and very careful control of the temperature if an improvement is to be seen and the strength of the metal is not to be affected. For these reasons, spot heating for fatigue improvements has been virtually discontinued.

Hammer peening with a round nosed tool or needle gun peening gives very good results although the noise produced may prevent their use. Shot peening can also be used to introduce compressive stresses at weld toes with equally good results. Compressive stresses can be induced in a component by over stressing - a pressure test of a pressure vessel is a good example of this - where local plastic deformation at stress raisers induces a compressive stress when the load is released. This technique needs to be approached with some care as it may cause permanent deformation and/or any defects to extend in an unstable manner resulting in failure.

Although the next techniques described are not as beneficial as hammer peening of the weld toes they have the advantage of being more consistent and easier to control. The techniques rely upon dressing the weld toes to improve the shape and remove the intrusion mentioned in article 79. The dressing may be carried out using a TIG or plasma-TIG torch which melts the region of the weld toe, providing a smooth blend between the weld face and the parent metal.

Alternatively the toe may be dressed by the careful use of a disc grinder but for best results the toe should be machined with a fine rotary burr as shown in Fig.3 and 4. Great care needs to be exercised to ensure that the operator does not remove too much metal and reduce the component below its minimum design thickness and that the machining marks are parallel to the axis of the main stress. Ideally the dressing should remove no more than 0.5mm depth of material, sufficient to give a smooth blend and remove the toe intrusion. The results of these improvement techniques are summarised in Fig.5.

Fig.3. Grinding tools

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Fig.4. Burr machining of weld toes

Fig.5. Improvement in fillet weld fatigue life

Whilst fatigue has resulted in some catastrophic and unexpected failures, the improvements in design life calculation methods, particularly the use of powerful software packages allowing detailed finite element analyses to be performed, has enabled engineers to approach the design of fatigue limited structures with far more confidence. This still means, however, that the designer has to recognise the effect of welds in the structure and must consider all possible sources of loading and ALL welds, even non-load carrying attachments that may be thought to be unimportant to service performance.

Fatigue FailureFatigue is a mechanism of failure experienced by materials under the action of a cyclicstress. It involves initiation and growth of a crack under an applied stress amplitudethat may lay well within the static capacity of the material. Discontinuities such aschanges in section or material flaws are favoured sites for fatigue initiation. Duringsubsequent propagation the crack grows a microscopic amount with each load cycle.The crack so-formed often remains tightly closed, and thus difficult to find by visualinspection during the majority of the life. If cracking remains undiscovered, there is arisk that it may spread across a significant portion of the load-bearing cross section,leading to final separation by fracture of the remaining ligament, or another failure modemay intervene such as jamming of a mechanism. Fatigue occurs in metals, plastics,composites and ceramics. It is the most common mode of failure in structural andmechanical engineering components. Fatigue failure is synonymous with the aviationindustry where square window frames within the initial design of the first commercial jetairliner the Comet 4 C caused fatigue failures and tragic loss of life on 2 full commercialaircraft at around 10,000 hrs of flight time before the fracture mechanism was fullyidentified and re-mediated and is the reason why we look out of oval windows whenever

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we should fly by jet aircraft.The phenomenon has been investigated extensively over many decades, particularly inmetals and alloys. As a result, design guidance is readily available in many texts and iswidely codified. Joints in materials are particularly susceptible to fatigue and thereforeneed to be designed with care for cyclic loading. Fatigue design rules for welded andbolted connections in steel can be found in many national standards, e.g. BS 7608 andBS 5400 widely used in the UK.MorphologyFatigue cracks generally exhibit a smooth surface and propagate at 90° to the directionof applied stress. The initiation points can usually be identified as weld flaws/features,machining marks or geometrical stress raisers. In some instances striations and beachmarks can be seen. Striations can be viewed using and electron microscope and arerecords of the crack growing under each loading cycle. Beach marks can be view withthe naked eye and can indicate a change in loading pattern. Both of these phenomenacan be used to estimate the fatigue crack growth rate. Fatigue cracks continue to growuntil the increasing level of stress cannot be supported with the final few cycles inducinglarger amounts of fracture surface and final fracture occurs.2The final fracture surface will show an area of fatigue failure emanating from the fractureinitiation point, with the fractured surface characterised by beach marks. These beachmarks may no longer be visible due to burnishing caused by metal/metal contact,though the final beach mark at the point of final failure is as a rule generally alwayspresent.

Striations (x1500)

Beach marks – initiation site arrowed

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Fatigue designThe standard method of representing fatigue test data is on an S-N curve. This plotseither the stress or strain range on the y-axis and the number of cycles to failure on thex-axis. The lower the stress range, the more cycles are required to cause failure. Whenpotted on logarithmic axes the data for a particular specimen type can be approximatedto a straight line between 105 and 107 cycles. Under constant amplitude loadingconditions most materials exhibit a fatigue limit. It is believed that tests performed atstress ranges below this limit will never cause a fatigue failure. For un-welded steels thefatigue limit occurs at approximately 2 million cycles, for welded steels and aluminiumalloys this is closer to 10 million cycles. Because of the relatively low fatigue limit,aircraft components made from aluminium alloys have a finite lifespan, after which theyare replaced. Fatigue is generally independent of rate of loading and temperatureexcept at very high temperatures when creep is likely. However, the presence of a3corrosive environment (e.g. sea-water) can have a significant detrimental effect onfatigue performance in the form of corrosion fatigue.

Flaw assessmentIn welded joints, fabrication flaws may give rise to premature fatigue failure, particularlyplanar flaws such as lack of fusion. Using fracture mechanics, the rate at which fatiguecracking will grow from such features can be estimated, and in this way tolerable flawsizes can be derived. British Standard 7910 provides detailed guidance on this methodof assessment.Factors to be considered when investigating a fatigue failureFatigue cracks initiate at areas of stress concentration such as discontinuities,weldments or sires of mechanical damage. They are a result of cyclic loading and canoccur at stress ranges well below the material’s UTS. It is of prime importance tounderstand the nature (vibration, thermal, mechanical etc.) and magnitude of theloading in order to prevent failure. Often the final failure of the component/structure willbe due to brittle or ductile fracture, therefore the fracture surface will show acombination of failure modes.RemediationFor weldments where fatigue is known to be a problem, life extension techniques such

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as weld toe burr machining, TIG dressing and peening can be used. These are effectivebut labour intensive and therefore expensive.4Brittle fractureBrittle fracture is the rapid run of a crack(s) through a stressed material. There is verylittle prior plastic deformation and so failures occur without warning. In brittle fracture thecracks run close to perpendicular to the applied stress, leaving a relatively flat surface atthe break. A brittle fracture surface may exhibit one or more of the following features.Some fractures have lines and ridges beginning at the origin of the crack and spreadingout across the crack surface. Others, some steels for example, have back-to-back Vshaped‘Chevron’ markings pointing to the origin of the crack. Amorphous materialssuch as ceramic glass have a shiny smooth fracture surface and very hard or finegrainedmaterials may show no special pattern.

Chevron fracture surface

Radiating ridge fracture surfaceIn common with fatigue fractures all brittle fractures require a point of initiation, andtherefore generally formed at areas of high stress concentration. This could be from aweld toe, undercut, arc strike, or could possibly be at the tip of a freshly initiated fatiguecrack, as is though to have been the case with the Liberty Vessels sunk during theSecond World War and which often sailed through the icy cold and tempestuous ArcticOcean in order to avoid detection and destruction from the German U Boat torpedoes.Fatigue cracks are though to have initiated at the square hatches through bad design,as in order to increase shipping production faster than shipping losses due to sinkingthe Liberty Vessels were the first welded vessels in the history of ship construction.5Ductile FractureWhen compared with brittle fractures, ductile fractures move relatively slowly and thefailure is usually accompanied by a large amount of plastic deformation. Ductile fracturesurfaces have larger necked regions and an overall rougher appearance than a brittlefracture surface. The failure of many ductile materials can be attributed to cup and conefracture. This form of ductile fracture occurs in stages that initiate after necking begins.

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Plane strain effectA condition in linear elastic fracture mechanics in which there is zero strain in a directionnormal to both the axis of applied tensile stress and the direction of crack growth. Underplane strain conditions, the plane of fracture instability is normal to the axis of principalstress. This condition is found in thick plates. Along the crack border stress conditionschange from plane strain in the body of the metal towards plane stress at the surface,this is displayed by the appearance of thin bands, caused by intense shear, that breakthrough to the free surface. The structure now becomes a mechanism, and whereplasticity breaks through to the surface shear lips will be observed.

Plane strain fracture: - plastic zone diameter ro much less than sample thicknessSynopsis1) Fatigue failuresGenerally produce beach marks indicating boundaries of plastic slip, generally > x 1 x 106cycles. The fracture initiation point forms generally from a stress concentration ie weldtoe, crack, or an abrupt change in section and can generally be identified at the epicentreof the beach mark/radii. Never the final, but very often the first mode of fracture, fatiguefailures are generally normal (90°) to the plain of the applied cyclic stress.2) Ductile failuresGenerally occur at 45° to plain of the applied stress with the fracture surface having arough or torn appearance. They may often occur as the second or final mode of failurein a fatigue specimen where the CSA can no longer support the load and are generallyaccompanied by shear lips. (Local plastic deformation)

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3) Brittle failuresGenerally occur at 90° to plane of the applied stress with the fracture surface having asmooth crystalline appearance. Again the fracture initiation point forms generally from astress concentration ie welded toe, crack, or abrupt change in section and can be oftenbe identified by the presence of chevrons, which point to the fracture initiation point.Failures that initiate as brittle fractures are unlikely to show evidence representing anyother forms of fracture morphology upon their surfaces.When in initiated as brittle fractures these surfaces do not show any plastic indicationsand if initiated as such will remain purely as brittle fractures, traveling in excess of thespeed of sound.4) Plane strain effectFlat areas occurring at 90° indicating plane strain effect may also appear centrallyupon fractured surfaces, and are caused by the inelastic behavior in larger materialthickness, in otherwise ductile specimens. It is thus possible to find a single fracturesurface showing 1 2 and 4 of the above characteristics, as in the ductile CTOD or cracktip opening displacement test shown below.

Creep and creep testing

Job Knowledge

The use of metals at high temperatures introduces the possibility of failure in service by a mechanism known as creep.

As the name suggests this is a slow failure mechanism that may occur in a material exposed for a protracted length of time to a load below its elastic limit (see Connect article No. 69), the material increasing in length in the direction of the applied stress. At ambient temperature with most materials this deformation is so slow that it is not significant, although the effect of low temperature creep can be seen in the lead on church roofs and in medieval glazing, where both materials have slumped under the force of gravity.

For most purposes such movements are of little or no importance. Increasing the temperature, however, increases the rate of deformation at the applied load and it is vitally important to know the speed of deformation at a given load and temperature if components are to be safely designed for high temperature service. Failure to be able to do this may result in, for example, the premature failure of a pressure vessel or the fouling of gas turbine blades on the turbine casing.

The drive for the more efficient use of fuels in applications such as power generation plant and gas turbines demands that components are designed for higher and higher operating temperatures, requiring new creep resistant alloys to be developed. To investigate these alloys and to produce the design data the creep test is used.

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In metals, creep failure occurs at the grain boundaries to give an intergranular fracture. Fig.1 illustrates the voids that form on the grain boundaries in the early stages of creep. The fracture appearance can be somewhat similar to a brittle fracture, with little deformation visible apart from a small amount of elongation in the direction of the applied stress.

Fig.1. The voids that form on the grain boundaries in the early stages of creep a)

b)

The creep test is conducted using a tensile specimen to which a constant stress is applied, often by the simple method of suspending weights from it. Surrounding the specimen is a thermostatically controlled furnace, the temperature being controlled by a thermocouple attached to the gauge length of the specimen, Fig.2. The extension of the specimen is measured by a very sensitive extensometer since the actual amount of deformation before failure may be only two or three per cent. The results of the test are then plotted on a graph of strain versus time to give a curve similar to that illustrated in Fig.3.

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Fig.2. Schematic of a creep test

Fig.3. Typical creep curve for steel

The test specimen design is based on a standard tensile specimen. It must be proportional (see Connect Article No. 69) in order that results can be compared and ideally should be machined to tighter tolerances than a standard tensile test piece. In particular the straightness of the specimen should be controlled to within some ½% of the diameter. A slightly bent specimen will introduce bending stresses that will seriously affect the results. The surface finish is also important - the specimen should be smooth, scratch free and not cold worked by the machining operation. The extensometer should be fitted on the gauge length and not to any of the other load carrying parts as it is difficult to separate any extension of these parts from that in the specimen.

Testing is generally carried out in air at atmospheric pressure. However, if it is necessary to produce creep data for materials that react with air these may be tested in a chamber containing an inert atmosphere such as argon or in a vacuum. If the material is to operate in an aggressive environment then the testing may need to be carried out in a controlled environment simulating service conditions.

Fig.3 shows that creep failure occurs in three distinct phases - a rapid increase in length known as primary creep where the creep rate decreases as the metal work hardens. This is followed by a period of almost constant creep rate, steady state or secondary creep and it is this period that forms the bulk of the creep life of a component. The third stage, tertiary creep, occurs when the creep life is almost exhausted, voids have formed in the material and the effective cross sectional area has been reduced. The creep rate accelerates as the stress per unit area increases until the specimen finally fails. A typical failed specimen is illustrated in Fig.4.

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Fig.4. Fractured test specimen

The creep test has the objective of precisely measuring the rate at which secondary or steady state creep occurs. Increasing the stress or temperature has the effect of increasing the slope of the line ie the amount of deformation in a given time increases. The results are presented as the amount of strain (deformation), generally expressed as a percentage, produced by applying a specified load for a specified time and temperature eg 1% strain in 100,000hrs at35N/mm 2 and 475°C.

This enables the designer to calculate how the component will change in shape during service and hence to specify its design creep life. This is of particular importance where dimensional control is crucial, in a gas turbine for instance, but of less importance where changes in shape do not significantly affect the operation of the component, perhaps a pressure vessel suspended from the top and which can expand downwards without being compromised.

There are therefore two additional variations on the creep test that use the same equipment and test specimen as the standard creep test and that are used to provide data for use by the designer in the latter case. These are the creep rupture test and the stress rupture test. As the names suggest both of these tests are continued until the specimen fails. In the creep rupture test the amount of creep that has occurred at the point of failure is recorded. The test results would be expressed as %age strain, time and temperature eg rupture occurs at 2% strain at 450°C in 85,000 hours. The stress rupture test gives the time to rupture at a given stress and temperature eg 45N/mm2 will cause failure at 450°C in 97,000 hrs. This data, if properly interpreted, is useful in specifying the design life of components when dimensional changes due to creep are not important since they give a measure of the load carrying capacity of a material as a function of time.

Relevant Specifications

BS EN 10291 Metallic Materials - Uniaxial Creep Testing in Tension.BS 3500 Methods for Creep and Rupture testing of Metals.ASTM E139 Conducting Creep, Creep Rupture and Stress Rupture Tests of Metallic Materials.BS EN ISO 899 Plastics - Determination of Creep Behaviour.

BS EN 761Creep Factor Determination of Glass- Reinforced Thermosetting Plastics- Dry Conditions.

BS EN 1225Creep Factor Determination of Glass- Reinforced Thermosetting Plastics- Wet Conditions.

Design - Part 1

Job Knowledge

This next series of Connect articles will look at welding design.

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Best practice in design is not simply a matter of deciding on the appropriate weld size or component thickness capable of carrying the service loads; there are many aspects of designing a welded component that need to be considered in addition to calculating permissible stresses. Weldability and mechanical properties such as tensile strength, toughness and fatigue resistance, all of which the designer must be familiar with, have been dealt with in a number of other Job Knowledge articles and will not be covered in this series on design.

In addition to selecting the material and specifying weld sizes, the designer must bear in mind that the decisions that he/she makes will directly affect the cost, safety and serviceability of the structure or component.

It is therefore necessary for the designer to:

select the most appropriate material select the most cost effective design of welded joint design the component to be welded by the most cost effective process specify the smallest weld acceptable for both service and fabrication use the smallest number of welds ensure that there is adequate access for both welding and inspection ensure that realistic dimensional tolerances are specified and can be achievedThe topics mentioned above involve a range of specialised technologies and it is therefore essential for the designer to seek advice from other professions such as metallurgists and welding engineers and not to rely solely upon their own judgement. This must be done before the design process has proceeded beyond the point of no return; sadly this is often not the case!

To begin let us look at some definitions. Firstly, the joint type or configuration of which there are five fundamental forms as shown in Fig.1. Note that there are no welds associated with these joint types.

a) In-line or butt joint

b) T-joint

c) Corner joint

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d) Lap joint

e) Edge joint

Fig.1. Joint types (a) - (e) 

These various joint types may be joined by only two weld types. Firstly, the butt weld where the weld is within the plane of the components being joined and secondly, the fillet weld where the weld is completely or mostly outside the plane of the components ( Fig.2). Plug and edge welds are somewhat special cases and will be discussed later.

a) Butt weld

b) Fillet weld

Fig.2. Weld types

A butt weld may be combined with a fillet weld to form a compound weld as illustrated in Fig.3:

a) Single-sided T-butt weld

b) Single-sided T-butt weld with superimposed fillet weld - a compound weld

Fig.3. Compound welds

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Fillet welds are probably the most common type of weld, particularly in structural steelwork applications, so this first section will look at some of the design considerations of fillet welds. They may be used to make T, lap and corner joints ( Fig.4).

a) T-joint fillet weld

b) Corner joint fillet weld

c) Lap joint fillet weld

Fig.4. Single-sided fillet welded joint types

A fillet weld is approximately triangular in shape, the size being defined by the weld throat or leg length as shown inFig.5.

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Fig.5. Terms used to describe features of a fillet weld

Fillet welds sizes should be specified preferably by referring to the throat thickness 'a' although the leg length 'z' is often used and can be easier to measure during weld inspection. Conventionally, the leg lengths are regarded as being of equal dimensions, the weld forming an isosceles triangle in cross section.

The convex fillet is generally undesirable for two main reasons. a)The junction of the weld metal with the parent metal at the weld toe can form a significant stress raiser and will adversely affect both fatigue life and brittle fracture resistance; b) the excess weld metal in the cap costs both time and money to deposit without contributing to joint strength. The concave fillet weld can be beneficial with respect to fatigue strength and, if required, the minimum throat thickness MUST be specified.

Fillet welds are less expensive to make than butt welds as there is no requirement to cut or machine a weld preparation. Although they are capable of carrying substantial loads they should not be used where the applied loads put the root of the weld in tension, particularly where the loading is dynamic - fatigue life in particular is drastically reduced. Where such loading is a possibility then a double sided T-joint should be made using two fillet welds ( Fig.6).

Fig.6. Preferred fillet welded joint type under bending loads

It is commonly thought that the fillet weld is an easier weld for the welder to make than a butt weld as the weld is deposited on solid metal. However, this is not necessarily the case when full fusion into the root of the weld is required. It is not unknown for highly skilled welders to fail a fillet weld qualification test where this is a design requirement. This is an important point and needs to be considered firstly by the designer asking if it is an essential requirement and secondly by the fabricator when pricing a contract.

This also raises the point that the fillet weld is extremely difficult to volumetrically examine using non-destructive testing techniques to confirm its internal soundness. This applies particularly to the root region where it is not possible to measure, with any degree of precision, any lack of fusion, slag entrapment etc. Therefore the same reliance on joint integrity, and hence service performance, should not be placed on a fillet weld as may be placed on a fully inspected butt weld.

Design Part 2

Job Knowledge

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The article in the last issue of Connect introduced the fillet weld, the least costly weld type to make since the components to be joined do not require flame cutting or machining of a weld preparation, the pieces can be propped against each other and the welder can then deposit a single pass of weld metal against a solid metal backing.

Whilst this sounds simple there are some aspects of making a fillet weld that must be taken into account (in addition to those already mentioned in the previous article Design part 1).

Cooling rates in a fillet weld are greater than in a similar thickness butt joint. There are three paths by which heat will be lost from the weld. This fact means that lack of fusion/cold start defects are more likely, particularly in high thermal conductivity metals such as aluminium and the risks of cold cracking are increased in carbon and low alloy steels. What may be acceptable in terms of heat input and/or preheat temperature for a butt weld may therefore not be acceptable with a fillet weld configuration. This point has sometimes been overlooked, particularly when welding on temporary attachments such as strongbacks, where quality control may be somewhat lax. This has led to major cracking problems for some fabricators.

Unlike a butt weld where the required weld throat is generally the thickness of the parent metal, the size of a fillet weld is determined by the loads that it is expected to carry. It can therefore be of any size that the designer specifies although there are practical limitations with respect to both minimum and maximum throat thickness.

With the conventional arc welding processes it is difficult to deposit a fillet weld with a throat less than some 2mm. This is in addition to the possibility of the lack of fusion/cold cracking mentioned above due to the rapid cooling rates experienced by small fillet welds. The maximum size of fillet weld is generally that of the thickness of the thinner of the two items being joined but very large fillet welds may cause unacceptable distortion and/or extremely high residual stresses. In addition, above a certain size it may be more economical to make a T-butt, rather than a fillet weld.

Although the throat thickness is regarded as being the most important dimension for design purposes it is a fact that mechanical failure of fillet welds is often along the fusion line or through the parent material itself. One reason for this in carbon or low alloy steels is that the weld metal is mostly substantially stronger than the parent metal.

As mentioned in Connect article No. 90 there are a variety of fillet weld shapes that make the accurate measurement of the throat thickness a little more difficult than may be first thought.

The throat is the shortest distance from the root to the face of the weld. To measure this dimension in a regular mitre or flat faced fillet weld is relatively simple. The shape is that of an isosceles triangle, the throat being 0.7 of the leg length. Convex, concave and deep penetration welds, however have throat thicknesses as illustrated in Fig.1.

a) mitre fillet weld

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c) concave fillet weld

b) convex fillet weld

d) deep penetration weld

Fig.1.Throat dimensions in fillet welds

It is apparent then, that measurement of either leg length or actual throat thickness alone is not reliable in determining the design throat thickness of a weld but that the weld shape must be taken into account. The excess weld metal of the convex weld gives no benefit with respect to design strength and, from a cost point of view, the fillet weld face should be as flat as possible.

The deep penetration weld is a very cost effective way of increasing the joint strength as only a proportion of the weld metal is from deposited filler metal. However, it is not possible to measure the throat of a deep penetration weld. To guarantee that the minimum design throat has been achieved it is necessary to control welding parameters and fit-up within very tight tolerances. This type of weld is therefore generally made using an automated or mechanised welding process (submerged arc or spray transfer MIG/MAG) in order to achieve sufficient and consistent control of the welding parameters.

When deciding on the size of a fillet weld it should be remembered that a small increase in throat thickness will result in a large increase in deposited weld metal as the cross sectional area of a fillet weld is a function of the square of the leg length (area = z2/2). Increasing the throat from, say, 5 to 6mm results in an increase in area and therfore weld metal of around 45%. This equates to almost 0.1kg extra weld metal per 1 metre length of weld. There are thus substantial cost and weight penalties to be paid if the joint is either over-specified by the designer or over-welded by the welder. There are no hard and fast rules about the point at

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which it is more economical to change from a fillet weld to either a double sided fillet weld or a partial penetration butt weld. Areas quoted in Fig.2 are worth bearing in mind when deciding on fillet weld sizes.

Fig.2. Relative cross sectional areas

For a fillet weld loaded in shear (the load parallel to the weld) the calculation of stress on the weld is simple; it is the load divided by the area of the weld throat.

Fig.3. Calculation of fillet weld throat

It is assumed for design purposes that fillet welds fail through the throat and it is therefore a simple matter to calculate the cross sectional area capable of carrying this applied load when the strength of the weld metal is known.

Note that the shear strength of a metal is generally around 70% to 80% of the tensile strength. This figure is often factored to give an acceptable margin of safety. In the UK for plain carbon steels a shear strength of 115N/mm2 is frequently used, enabling the throat thickness to be calculated from the simple formula throat:- throat 'a' = P/(L x 115).

The throat dimensions of a double fillet weld T-joint loaded in tension can be determined using the same approach. Note, however, that this is a very simplistic calculation and does not take into account any other stresses (bending, torsion etc) that the weld may experience. It is however beyond the scope of these brief articles to cover in any depth the stress analysis of welds.

Design Part 3

Job Knowledge

Part 1 Part 2 Part 4 Part 5

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Fillet welds may be combined with full or partial penetration butt welds - a combination weld. The designer is therefore required to decide whether to use a T-butt weld, a fillet weld or a combination of the two. In making this decision cost is a major factor.

As mentioned in Job knowledge 91, the fillet weld requires no weld preparation, is easy to deposit and is often regarded as the cheapest weld of all to make. However cross sectional area, and therefore cost, increases as a function of the square of the leg length. Assuming the same strength requirements from the fillet welds as for the T-butt welds it becomes more economical to use a double sided full penetration T-butt joint at a plate thickness of around 30mm.The accuracy of this figure should be treated with caution as it is dependent on many factors such as the weld preparation costs and included angle.

Welding position is an additional factor. It may be more economical to deposit a butt weld in the flat position, where large diameter electrodes and high welding currents can be used, rather than a double sided fillet weld where one weld must be made in the overhead position ( Fig.1).

Fig.1. Flat position T-butt weld vs overhead fillet weld

An additional benefit from using a T-butt weld is that this weld type provides a direct transfer of force through the joint, giving a better performance under fatigue loads. Many design specifications will also have lower allowable stresses for a fillet weld compared with a butt weld and this can have a significant effect on cost, particularly when designing to match the strength of thicker plates.

It should be remembered that it is difficult, if not impossible, to examine a fillet weld volumetrically using radiographic or ultrasonic techniques and the internal weld quality is therefore entirely dependent on the skill and integrity of the welder. The comments on T-joints also apply to corner joints where two fillet welds may be more economical than one large fillet as shown in Fig.2. However, remember that one weld may need to be made in the overhead position if the component cannot be turned.

Fig.2. Corner Joints: Area of weld in a) -50mm2; and b) -25mm2

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From the foregoing it is obvious that the decision to use fillet welds, T-butt joints or combination welds is not as straightforward as it may first appear and there are numerous factors that must be taken into account.

Butt joints are those welds where the weld metal is contained within the planes of the surfaces of the items being joined. The weld throat may be the full section thickness, a full penetration joint, or a proportion only - a partial penetration joint. Welds may be 'single sided joints', welded all from one side, or 'double sided', welded from both sides, ( Fig.3).

Fig.3. Full and partial penetration welds

Except for very thin plate, arc welded butt joints require a weld preparation to be flame cut or machined along the joint line. The conventional arc welding processes can penetrate into the base metal by only a limited amount. The maximum penetration in conventional TIG or manual metal arc (SMAW) welds is in the region of 3mm, MAG (GMAW) welds around 6mm and submerged arc some 15mm.

In order to weld the full thickness of a plate and achieve the weld throat thickness required by design it is therefore necessary to cut away sufficient metal along the joint line so that the welding electrode has access to the root of the joint, enabling the root pass to be deposited and then the remainder filled to complete the joint. A weld preparation, the 'weld prep', is therefore formed along the joint line using flame cutting, plasma cutting or machining. Figure 4identifies the key features of a 'single bevel' weld preparation and those of a 'single-V' joint.

The smaller the included angle, the less access this will give to the root and the greater is the risk of defects such as lack of side wall fusion. This reduced access may, however, be compensated for by an increase in the root gap.

The bevel angles and the root gap will depend upon the process(es) used to make the joint and the material thickness. A narrow included angle requires less weld metal and therefore is more economical as the thickness increases. A downside to this is that the narrower the angle the more difficult access becomes and the risk of welding defects as mentioned above.

Too wide a root gap will result in a loss of control of the weld pool and melt through giving an irregular and excessive penetration bead. This may be overcome by using a backing strip if this is permitted by the service conditions.

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Fig.4. Single bevel weld preparation

The choice of the weld preparation is therefore a compromise between maintaining adequate access and minimising the weld volume.

If a high quality root bead is required and access is not available to the root side of the weld e.g. in a pipe carrying fluids or in high pressure service, then an acceptable condition can be achieved using the TIG process to make the root bead. A typical pipe butt weld set-up would be 60° included angle, 1mm to 2mm root gap and a zero to 1.5mm thick root face.

Where access to the reverse side of the joint is available, the condition of the penetration bead is less important as the root bead can be ground to sound metal and a sealing pass deposited.

A reduction in weld volume can be achieved by the use of a 'J' preparation as shown in Fig.5. This preparation, unlike the straight chamfer of the 'V' preparation which can be flame cut, must be machined.

Fig.5. Key features of single sided 'J' preparation

This can be an expensive operation, which is why this type of weld is used only on thick joints, where the saving in deposited weld metal outweighs the cost of machining, or where very high quality root beads are required.

Machining of the weld preparation dictates that the dimensions, particularly that of the root face thickness, can be controlled far more closely than is possible with flame cutting and therefore a more accurate fit-up can be achieved.

It is often used on orbitally TIG welded pipe butt joints where a machined joint enables the tolerances required by a fully automatic process to be achieved.

Design Part 4

Job Knowledge

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Part 1 Part 2 Part 3 Part 5

The previous article looked briefly at butt weld design where mention was made of the increased risk of producing defects as the bevel angle or the root gap is reduced. Bevelling the plate edges allows access to all parts of the joint, enabling good fusion throughout the weld to be achieved. The bevel can be on one or both edges of the items to be joined. What is important is the included angle which is dictated by the need both to achieve the correct torch/electrode angle and to maintain the required arc length and wire stick-out. as shown in Fig.1. The angle on a single bevel joint, as in Fig.1(c) obviously needs to be greater than that on a double bevel V-joint if access problems are not to be encountered. Experience has shown that a weld preparation angle of 45° on a single bevel joint is usually sufficient to allow adequate access.

Fig.1. Effect of a narrow weld preparation angle

A similar effect is produced by too narrow a root gap where, as above, there is insufficient access to permit a correct arc length to be achieved and the arc cannot be placed in the correct position. Conversely, too wide a root gap on an unbacked weld will require a large, wide weld pool to bridge the gap, resulting in melt through, a loss of control of the pool and the formation of localised excess weld metal - known colloquially as 'noddies' or 'dangleberries'.

As may be guessed from the above, the most problematic region in a weld is that of the root pass. Single sided joints require dimensionally accurate weld preparations and fit-up and skilled welders to ensure full penetration welds with an acceptable root contour. The best root pass appearance using conventional arc welding processes will be achieved using the TIG process but acceptable root conditions can also be achieved with MMA, MIG/MAG and FCAW welding.

When welding, it is obviously easier for the welder when there is a sound base on which to deposit the weld metal; hence the need for a very skilled welder when making full penetration single sided welds.

When permitted by design, it is possible to use partial penetration welds as illustrated in Fig.3 of Connect article no. 92. However, note that this type of joint is not recommended when fatigue is an issue.

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Where access to the reverse side of a partial penetration weld is available, then the fabricator has the option of depositing a sealing pass. Remember, however, that most welding processes have only limited penetration and there is a real risk that not all of the unfused land will be melted away. To be certain of removing the unfused land, 'back gouging' the root and filling the groove with sound weld metal is generally carried out. Backgouging, or removal of the unfused land, can be done by any of the conventional metal removal techniques; machining, arc air gouging, chipping, grinding etc (Fig.2). Of these methods, arc air gouging is probably the most cost effective and can produce a smooth contoured U-shaped groove with an included angle of 50 to 60 degrees, allowing adequate access.

The back gouging must be sufficiently deep that any lack of penetration is removed. To confirm that this has been done it is good practice to perform magnetic particle or liquid penetrant inspection of the gouged groove.

Fig.2. Backgouging to achieve full penetration

An alternative to backgouging, or when access to the reverse side is not available, is to use a backing strip which will provide support for a fully penetrated root pass. The backing strip may be permanent or temporary, (see Fig.3 below).

Fig.3. Various forms of backing

The permanent backing strip weld does not have as good a performance in fatigue loading as a single sided TIG root butt weld and the crevice is a site for preferential corrosion. Whether a permanent backing strip weld is acceptable for service is therefore a design decision.

In addition to providing support for the root pass, a further major advantage of the backing strip weld is that fit-up tolerances may be relaxed as the strip acts as a locating feature. This is particularly so when pipe butt welding where the strip forms a spigot on which to centre the joining pipe. In addition root gap may be varied, the only real limitations being those of cost; the wider the root gap the greater the volume of weld metal and distortion.

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The strip must be compatible with the filler metal and the parent base metal. It must be correctly fitted, in close contact with both edges of the weld preparation and welded into position using intermittent tack welds. Any gap between the backing strip and the plate edges is a site for slag entrapment and results in a poor root profile. To ensure full fusion in the root of the weld it is advisable to use a feather edge and to direct the welding arc at the plate/pipe edges.

When a permanent backing strip cannot be used, then a temporary backing bar may be used (conventionally a permanent backing is known as a 'strip', a temporary backing as a 'bar'). As the name suggests this is a backing that is easily removed at the end of the welding operation; it has not become fused to the root pass.

It may be made of a ceramic or of copper, chromium plated for use on stainless steel and nickel based alloys to prevent contamination. Austenitic stainless steel has also been used. The metal backing bars may be water cooled to aid heat loss and may be grooved to provide a mould for the molten weld metal. Welding conditions and fit-up must be carefully controlled to prevent the welding arc from impinging directly on the bar, otherwise there maybe melting of the bar and contamination of the weld pool.

Ceramic backing bars can be obtained in a variety of sizes with shaped grooves to form a weld pool mould. They may be rigid bars of ceramic or articulated such that they can be wrapped around the inside diameter of a pipe or tube. Ceramic tapes are also available, as illustrated above.

These tapes have wide strips of adhesive either side of the ceramic tile to enable the tape to be held in place during welding and peeled off on completion. As with the permanent backing strip, care needs to be taken to ensure that the ceramic tile is in close contact with the metal surfaces otherwise slag and/or weld metal will run into the gap, giving an irregular weld root.

Design Part 5

Job Knowledge

Part 1 Part 2 Part 3 Part 4

The previous Job Knowledge articles looked at fillet and partial/full penetration butt welds. The final three weld types to be dealt with in this series on weld design are the edge weld, the spot weld and the plug weld.

The edge weld is a specialised weld that has limited fields of application and is mostly used for the joining of sheet metal components although it may be used for the fabrication of tube to tubesheet welds. The edge weld is frequently used as an alternative to a corner weld where achieving an accurate fit may be difficult, particularly on thin section components. Instead, by raising a flange on one of the components and clamping the two components together a weld can be made along the edge. Sealing the lid on a can is one ideal application as the lid can be pushed in to the can, resulting in a minimal gap and a self jigged joint (Fig.1). The weld size and penetration is limited so this weld type is generally only possible on thin components using methods such as TIG, plasma TIG or the power beam welding processes.

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Fig.1. Edge weld used to seal container lid

This type of edge weld may also be used for tube to tubesheet welding where, by machining a pintle onto the tubesheet, the tube can be inserted through the tube hole and an edge weld made, (Fig.2) This has the advantage that the heat sink is more evenly balanced when attempting to weld a thin tube to a thick tubesheet. In tubesheets of limited weldability or where postweld heat treatment is essential it is possible to deposit a ring of weld metal round the tube hole. This ring may then be machined to provide the pintle so that the residual stresses are reduced and the tube/tubesheet weld is made in good weldability weld metal. This results in a reduction in residual stress in the tubesheet and a reduction in the risk of cracking.

Fig.2. Edge weld used to weld tube to tubesheet joints

Alternatively, if PWHT is required the tubesheet and its weld rings can be PWHT'd, the pintles machined on and non-destructively examined (NDE) and the tube/tubesheet welds made in the thin section, removing the need for a second PWHT cycle. Because of the accuracy of these machined joints the welding process, generally TIG, is frequently mechanised or fully automated.

The spot weld, Fig.3, is normally associated with resistance welding where two thin sheets are overlapped and held in close contact by pressure from the welding electrodes during the welding cycle. The resistance spot weld could therefore be regarded as self jigging. Spot welding with the arc welding processes also uses a lap type joint but presents a more difficult problem in that the joint must be firmly clamped together such that there is no gap between the two surfaces. Failure to do this means that the weld metal may spill into the gap and full fusion to the underlying plate may not be achieved. Good jigging and fixturing is therefore essential.

Fig.3. Spot welds

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Applications of this joining method include sheet metal work and the lining ('wallpapering') of ducts, tanks etc with thin, corrosion-resistant sheets. The greatest strength of the welds is developed when the welds are in shear parallel to the plate surfaces.

As mentioned earlier, penetration into the parent metal from the various arc welding processes is limited, around 4mm with TIG (perhaps as much as 10mm with activated flux TIG), 10mm with plasma-TIG and 6mm with MAG welding. The thickness of the upper plate that must be fully penetrated to provide a sound weld is therefore similarly limited. An additional problem with MAG welding is that the filler wire is fed continuously into the weld pool so that a large lump of excess weld metal may be deposited on the plate surface. Autogenous TIG or plasma-TIG will give a weld flush with or slightly below the plate surface. The process can be partially mechanized. Special torches are available that, when held against the plate surface, give the correct electrode/work piece distance and timers on the welding power source that may be set to give the desired arc time.

To enable thicker plate to be joined by 'spot welding' a circular or elongated hole may be machined through the top plate, enabling either a plug or a slot weld to be made by filling the hole with weld metal. Whilst this may seem tobe a simple and easy process the strength of this type of joint depends upon full fusion of the weld metal with the vertical wall of the hole cut into the upper plate, see Fig.4. As with a fillet weld, lack of fusion in this area will result in a reduction in the throat thickness of the joint. It is therefore essential that the welder directs the welding arc into the bottom corner of the joint and does not simply puddle the weld metal into the hole. With small diameter plug welds this can be a difficult and skilled operation and welders need to be adequately trained to ensure that they can achieve full fusion.

Fig.4. Plug and slot welds

Since the strength of the plug or slot weld is determined by the throat it may not be necessary to fill the hole completely unless the weld must be flush with the surface of the plate for cosmetic reasons. Besides being unnecessary from the point of view of joint strength, a completely filled hole will have high residual stresses. These may cause unacceptable distortion and will increase the risk of cold cracking in carbon and low alloy steels.

This brief series of Job Knowledge articles has concentrated on the design of joints for welding. The designer also needs to remember that, not only must the joints be suitable for welding, they must in addition enable any non-destructive testing required by the contract or specification to be carried out. Provision therefore needs to be made to allow adequate access for the positioning of radiographic film and the radiation source, or to enable the correct scanning patterns to be used if the joint is to be ultrasonically tested.

Whilst NDE of butt welds is reasonably straightforward, radiography or ultrasonic examination of fillet welds is not generally regarded as being possible. The designer must therefore take into account the possibility of undetected defects in this type of joint.

Calculating weld volume and weight

Job Knowledge

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Calculating the volume of a weld is one of the first steps to be taken when estimating the cost of making a weld.

With this information, and knowing the deposition rate of the process, it is possible to determine the arc time (the length of time that an arc is burning and depositing weld metal) and the amount of welding consumables required to fill the joint. Both of these are required in order to calculate the cost of making the weld. Costing will be dealt with in future Job Knowledge articles.

Determining the volume of a weld requires some knowledge of basic geometrical calculations to determine the area of the weld and multiply this figure by its length. The first step then is to calculate the cross sectional area of the joint.

With a fillet weld or a 45° single bevel joint this is relatively simple but the calculations become lengthier as the weld preparation becomes more complex. Fig.1 illustrates how simple this calculation is for an equal leg length fillet weld; the area of such a weld is half the square of the leg length, Z. When using this formula do not forget that welders seldom deposit precisely the size of weld called up on the drawing or in the welding procedure and that there may be some excess weld metal on the face of the weld.

Fig.1. Area of an equal leg length fillet weld

An asymmetrical fillet weld is a little more difficult; the area of a triangle is given by the base Z2 times the height Z1divided by 2 so when a fillet weld is deposited with unequal leg lengths the area can be calculated from multiplying the throat, a, by the length of the face I and divided by 2 as illustrated in Fig.2.

Fig.2. Area of an unequal leg length fillet

Turning now to butt welds, the calculations become a little more complex.

There are three factors that determine the volume of the weld in a single V butt weld. These are the angle of the bevel, b, the excess weld metal and the root gap, g, as illustrated in Fig 3. To calculate the area of this weld we need to be able to add together the areas of the four components illustrated in Fig.3.

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Fig.3. The four areas of a single-V butt weld

The dimension 'c' is given by (tan b x t); the area of a single red triangle is therefore t(tan b x t)/2. The total area of the two red regions added together can be calculated using the formula 2t(tan b x t)/2 or t(tan b x t).

The width of the weld cap, w, is given by W = 2(tan b x t) + g.

The area of the excess weld metal is approximated by the formula (W x h)/2.

The area provided by the root gap by g x t.

The bevel angles, b, most often used are 10° = (tan 0.176), 15° = (tan 0.268), 22.5° = (tan 0.414) 32.5° = (tan 0.637) and 45° = (tan 1.00). As will become obvious when the weight is calculated, it is easier to ensure that the decimal point is in the right place if centimetres are used in the calculations rather than millimetres.

As a worked example, if the weld is in a plate 2.5cm thickness, 0.3cm root gap, 65° included angle (b = 32.50°; tan 32.5° = 0.637) and with a cap height of 0.2cm we have:-

1. c = tan32.5 x 2.5 = 0.637 x 2.5 = 1.59cm2. w = 2(0.637x2.5) + 0.3 = 3.485cm so the area of the cap = (3.485x0.2)/2 = 0.348 sq. cm.

3. area of the orange area = 0.3 x 2.5 = 0.75 sq.cm.

4. area of the two red areas = 2 x (1.59 x 2.5)/2 = 3.97sq.cm.

This gives a total area of 5.07sq cm. The volume can then be calculated by multiplying the length of the weld by the area - ensuring that this length is also given in centimetres!

Conventionally, the volume is often expressed in cubic centimetres (cu.cm). per metre so in this example the volume is 507 cu. cm/metre.

To obtain the weight of weld metal this figure is then multiplied by the density of the alloy. Table 1 gives the density of some of the more common alloys in gm/cu.cm. Note that with some alloys the alloying elements can change the density quite significantly.

Table 1. Densities of some of the more common alloys.

Alloy Density (gm/cm3)iron 7.870.25% carbon steel 7.8612%Cr steel 7.70304 stainless steel 7.92

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nickel 8.9080/20 Ni.Cr 8.40625 type alloy 8.44copper 8.9470/30 brass 8.537% Al bronze 7.89aluminium 2.70Al 5052 2.65Al 7075 2.8The weight of weld metal to fill one metre length of the joint described above would therefore be; in carbon steel (507 x 7.86) = 3985gms or 3.98kgs/metre; in a 5XXX series aluminium alloy (507 x 2.65) = 1343gms, 1.34kgs/metre.

Calculating the weight of weld metal in double sided V-joints uses the same approach by dividing the weld into its individual 'V's and adding the products.

A J-preparation, however, adds another area into the equation; that of the half circle at the root of the weld, see Fig.4. The formulae given above to calculate 'c', the area of the two red components and the excess weld metal remain unchanged but the width of the cap must be increased by 2r. There are also the two areas, 'A' and 'B', to calculate and the two white root radius areas to be added to the total.

Fig.4. Single 'U' preparation (other notation as in Fig.3)

The relevant formulae are thus:

1. the dimension 'c' is given by (tan b x (t-r)); the total area of the two red regions is therefore given by the formula 2((t-r)(tan b x (t-r))/2 or ((t.-r)(tan b x (t-r)).

2. the width of the weld cap, w, is given by w = 2(tan b x (t-r)) + g +2r.

3. the area of the excess weld metal is given by the formula (w x h)/2.

4. the area 'A' is (t-r) x (2r +g).

5. the area 'B' is g x r.

6. the root radius area is (πr2)/4

For a double-U preparation it is necessary to calculate the areas of both sides and add these together.

Having calculated the weight of weld metal required to fill a weld preparation it is then possible to calculate the weight of filler metal required (these two figures are not necessarily the same) and to estimate the time

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required to deposit this weld metal; both essential in order to arrive at a cost of fabricating the weld. This will be covered in future Job Knowledge articles.

Welding consumables - Part 1

Job Knowledge

Part 2 Part 3 Part 4 Part 5

The next series of articles will cover welding electrodes and filler metals, beginning with a brief look at the requirements for a flux. Whether a flux is in an electrode coating or is in granular form, as in a submerged arc flux, the requirements are the same.

The flux must be capable of providing a protective shield to prevent atmospheric contamination of the electrode tip, the filler metal as it is transferred across the arc and the molten weld pool. Generally, it does this by decomposing in the heat of the arc to form a protective gaseous shield.

It must be capable of removing any oxide film (failure to do so will result in lack of fusion defects and oxide entrapment). It does this by reacting chemically with the oxide.

It should improve mechanical properties by providing clean, high quality weld metal and perhaps by transferring alloying elements across the arc.

It must be capable of providing the desired weld metal composition, again by transferring alloying elements across the arc.

It should aid arc striking and arc stability. It should produce a slag that will shape the molten pool and hold the pool in place during positional

welding if required. Any slag should be readily removable and preferably self-detaching. It should not produce large amounts of fume and any that it does should not be harmful to the

welder.These requirements have resulted in a multitude of different consumables, many being apparently identical, and this can make filler metal selection a difficult and confusing task. This article attempts to give some insight into the various types of flux coated manual metal arc (MMA) electrodes before moving on in later articles to look at other types of welding fluxes.

Most MMA electrodes can be conveniently divided into three groups by their coating composition. These are cellulosic, rutile and basic coatings, each of which gives the electrode a distinctive set of characteristics.

Cellulosic electrodes contain a large proportion of cellulose, over 30% and generally in the form of wood flour. This is mixed with rutile (titanium dioxide, TiO2 ), manganese oxide and ferro-manganese and is bonded onto the core wire with sodium or potassium silicate. Moisture content of these electrodes is quite high, typically 4 to 5%. The cellulose burns in the arc to form a gas shield of carbon monoxide, carbon

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dioxide and, in conjunction with the moisture in the coating, produces a large amount of hydrogen, typically 30 to 45ml hydrogen/100gm weld metal.

The hydrogen raises arc voltage and gives the electrodes their characteristics of deep penetration and high deposition rates. The high voltage requires a high open circuit voltage of around 70 volts to allow easy arc striking and to maintain arc stability. The forceful arc also results in appreciable amounts of weld spatter and this limits the maximum current that can be used on the larger diameter electrodes. A thin, friable and easily removed slag is produced, giving a rather coarsely rippled weld profile. The slag is also fast freezing so that, unlike most other electrodes, they can be used in the vertical down position - 'stove piping'.

Electrodes with a sodium silicate binder can be used only on DC electrode positive (reverse polarity). Those with a potassium silicate binder can be used either DC electrode positive or on AC. The electrodes require some moisture in the coating to aid the running characteristics and they must never be baked, as may be done on basic coated electrodes. This has the advantage that the electrodes are tolerant to site conditions. If they become damp, drying at a temperature of around 120°C will be sufficient.

Electrode compositions are only available for welding low carbon non-alloyed steels although nickel additions may be made to improve notch toughness. Charpy-V values of around 27J at -20°C are possible in the unalloyed electrodes. The high hydrogen level means that any steel welded with these electrodes should be selected to have a very high resistance to hydrogen induced, cold cracking (see Connect articles numbers 45 and 46). They should not be used without giving due consideration to the steel composition, restraint and the need for preheat. The characteristics of deep penetration, high deposition rates and the ability to be used vertically down means that the main use for these electrodes is for cross country pipelining although they are used to a more limited extent for welding storage tanks.

Rutile coatings, as the name suggests, contain a large amount of rutile, titanium dioxide, typically around 50%, in addition to cellulose, limestone (calcium carbonate), silica (SiO2) mica (potassium aluminium silicate), ferro-manganese and some moisture, around 1 to 2%. Binders are either sodium or potassium silicate. The cellulose and the limestone decompose in the arc to form a gas shield containing hydrogen (around 20ml/100gm weld metal) carbon monoxide and carbon dioxide. The electrodes have medium penetration characteristics, a soft, quiet but stable arc and very little spatter, making them a 'welder friendly' electrode. Striking and re-striking is easy and the electrodes will run on very low open circuit voltages. The electrodes produce a dense covering of slag that is easily removed and gives a smooth evenly rippled weld profile.

The presence of cellulose and moisture means that the electrodes produce relatively high levels of hydrogen, perhaps 20 to 25ml/100gm weld metal. This restricts their use to mild steels less than 25mm thickness and thin section low alloy steels of the C/Mo and 1Cr1/2Mo type. Mechanical properties are good and Charpy-V notch toughnesses of 40J at -20°C are possible. They are probably the most widely used general purpose

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electrode. Rutile coated austenitic stainless steel electrodes can be obtained and can be used in all thicknesses as cold cracking is not a problem with these alloys.

Rutile electrodes, like cellulosic electrodes, require some moisture in the coating and they should not be baked. If they become damp, re-drying at around 120°C should be sufficient. Those electrodes with a sodium silicate binder can be used on DC electrode negative or AC. Electrodes with the potassium silicate binder can be used on both polarities and on AC. The potassium silicate binder electrodes generally have better arc striking and stability characteristics than the sodium silicate binder types and a more readily detachable slag.

The next article will look at the basic, low hydrogen electrodes and some of the other less common types of coatings.

Welding consumables - Part 2

Job Knowledge

Part 1 Part 3   Part 4   Part 5  

The previous article, Part 1, dealt with the cellulosic and rutile electrodes. This article will cover the basic, iron powder and acid electrodes.

The description 'basic' originates from the chemical composition of the flux coating which contains up to perhaps 50% of limestone, calcium carbonate (CaCO3). This decomposes in the arc to form a gas shield of carbon monoxide/dioxide.

In addition to the limestone there may be up to 30% of calcium fluoride (CaF2) added to lower the melting point of the limestone and to reduce its oxidising effect. Also deoxidants such as ferro-manganese, ferro-silicon and ferro-titanium are added to provide de-oxidation of the weld pool.

Other alloying elements such as ferro-chromium, ferro-molybdenum or ferro-nickel may be added to provide an alloy steel deposit. Binders may be sodium silicate, only for use on DC+ve current, or potassium silicate which enables the electrodes to operate on both direct and alternating current.

The gas shield from basic electrodes is not as efficient as that from the rutile or cellulosic types and it is necessary to maintain a constant short arc if porosity from atmospheric contamination is not to be a problem. The electrodes are particularly sensitive to start porosity because of the length of time taken to establish an efficient protective shield. An essential part of welder training is familiarisation with the technique of starting the weld ahead of the required start position and moving back before proceeding in the direction of welding.

The penetration characteristics of basic electrodes are similar to those of rutile electrodes although the surface finish is not as good. The slag cover is heavier than rutile electrodes but is easily controlled, enabling the electrodes to be used in all positions. High limestone coatings have been developed that enable a limited range of electrodes to be used in the vertical-down (PG) position. The weld pool blends smoothly into the parent metal and undercutting should not occur.

The slag is not as easily removed as with rutile or cellulosic electrodes but the low melting point means that slag entrapment is less likely. The chemical action of the basic slag also provides very clean, high quality weld metal with mechanical properties, particularly notch toughness, better than that provided by the other

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electrode types. A further feature of these electrodes is that the welds are more resistant to solidification cracking, tolerating higher levels of sulphur than a rutile or cellulosic electrode. This makes them valuable if it becomes necessary to weld free cutting steels.

The basic electrode is also known as a low hydrogen rod ('lo-hi'). The coating contains no cellulose and little or no moisture provided the electrodes are correctly handled. When exposed to the atmosphere, moisture pick-up can berapid. However, baking the electrodes at the manufacturers' recommended baking temperature, generally around 400°C, will drive off any moisture and should provide hydrogen levels of less than 5ml/100g weld metal. After baking the electrodes need to be carefully stored in a holding oven at a temperature of some 120°C to prevent moisture pick-up.

Many manufacturers now provide electrodes in hermetically sealed vacuum packs with hydrogen levels guaranteed to be less than 5ml/100g weld metal. These are particularly useful in site applications where there is a need to maintain very low hydrogen levels and baking and storage facilities are not available. The electrodes are taken directly from the pack and can be used for up to 12 hours from opening before sufficient moisture has been absorbed to require baking.

Basic, low hydrogen electrodes are therefore widely used in a variety of applications where clean weld metal and good mechanical properties are required. They can be obtained with alloyed core wires and/or ferro-alloy additions to the coating to give very wide selection of weld metal compositions, ranging from conventional carbon steels, creep resistant and cryogenic steels and duplex and stainless steels. Where high quality, radiographically or ultrasonically clean weld metal is a requirement, such as on offshore structures and pressure vessels, basic electrodes will be used.

Developments over the last 20 or so years have enabled carbon-manganese steel consumables to give good Charpy-V and CTOD values at temperatures down to -50oC. The low hydrogen capabilities also mean that basic electrodes would be used for the welding of thick section carbon steels and high strength, high carbon and low alloy steels where cold cracking is a risk (see Job knowledge articles Nos. 45 and 46).

In addition to the 'standard' cellulosic, rutile and basic electrodes discussed above, electrodes may be classified as 'high recovery'.

By adding substantial amounts of iron powder, up to 50% of the weight of the flux coating, to either basic and rutile electrode coatings it is possible to deposit a greater weight of weld metal than is contained in the core wire. These electrodes are described as having an efficiency above 100% eg 120%, 140% etc and this 3 digit figure is often included in the electrode classification.

The electrodes have thicker coatings than the 'standard' electrodes which can make them difficult to use in restricted access conditions. They are, however, welder friendly with good running characteristics and a smooth stable arc. The iron powder not only melts in the heat of the arc to increase deposition rate but also enables the electrode to carry a higher welding current than a 'standard' electrode.

The iron powder is electrically conducting, so allowing some of the welding current to pass through the coating. High welding currents can therefore be used without the risk of the core wire overheating, thus increasing both the burn-off and the deposition rates. The high recovery electrodes are ideally suited for fillet welding, giving a smooth, finely rippled surface with a smooth blend at the weld toes. They are generally more tolerant to variations in fit-up and their stability on low open circuit voltages means that they are very good at bridging wide gaps. However, the large weld pool means that they are not suited to positional welding and are generally confined to welding in the flat (PA) and horizontal-vertical (PC) positions.

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The last type of electrode covering is described as 'acid'. These electrodes have large amounts of iron oxides in the flux coating which would result in a high oxygen content in the weld metal and poor mechanical properties. It is therefore necessary to incorporate large amounts of de-oxidants such as ferro-manganese and ferro-silicon in the flux. Although they produce smooth flat weld beads of good appearance and can be used on rusty and scaled steel items the mechanical properties tend to be inferior to the rutile and basic coated electrodes. They are also more sensitive to solidification cracking and are therefore little used.

Welding consumables - Part 3

Job Knowledge

Part 1 Part 2 Part 4 Part 5

The last two articles covered the various types of manual metal arc consumables that are available.

In order to be able to specify the type of flux coating, welding characteristics and chemical composition of an electrode for a particular application, there needs to be some standardised method of unique identification that is universally recognised.

This requirement has led to the writing of a series of consumable specifications that enable an electrode to be easily and uniquely identified by assigning a consumable a 'classification'. The two MMA electrode classification schemes that will be dealt with in this month's article are the EN (Euronorm) and the AWS (American Welding Society) specifications. There is insufficient space to cover in detail the whole range of compositions for MMA electrodes so the emphasis here will be on the carbon steel filler metals.

The European specification for non-alloy and fine grained steel MMA electrodes is EN 499. This divides the classification or designation number into two parts. Part 1 is a compulsory section that requires symbols for the process, strength and elongation, impact strength, the chemical composition and the type of flux coating. The second part is optional and includes that includes symbols for the type of current and metal recovery, the welding position(s) that the electrode can be used in and for the maximum hydrogen content of the deposited weld metal (NOT the electrode).

The designation of a covered electrode begins with the letter 'E'. This tells us that this is a covered electrode intended for MMA welding. The next two numbers give the minimum yield strength that may be expected as shown in Table 1.

Table 1 Strength and elongation symbols

Symbol Min Yield StrengthN/mm

Tensile StrengthN/mm

MinimumElongation %

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35 355 440 - 570 2238 380 470 - 600 2042 420 500 - 640 2046 460 530 - 680 2050 500 560 - 720 18The next symbol indicates the temperature at which an average impact value of 47J can be achieved, as shown inTable 2.

Table 2 Impact value symbol

Symbol Temperature foraverage of 47J °C

Z No requirementA +200 02 -203 -304 -405 -506 -60The third mandatory symbol is for the composition. Although the specification title (non-alloy and fine grained steels) suggests that the electrodes have no alloying elements present, up to 3% Ni and NiMo electrodes are included, seeTable 3. (This symbol is only applied where the electrode contains ≥0.3Mo or ≥0.6Ni).

Table 3 Chemical composition symbols

Symbol Chemical composition % max or range  Mn Mo Ni

No symbol 2.0 - -Mo 1.4 0.3 - 0.6 -MnMo >1.4 - 2.0 0.3 - 0.6 -1Ni 1.4 - 0.6 - 1.22Ni 1.4 - 1.8 - 2.63Ni 1.4 - >2.6 - 3.8Mn1Ni >1.4 - 2.0 - 0.6 - 1.21NiMo 1.4 0.3 - 0.6 0.6 - 1.2Z Any other agreed compositionThe fourth symbol indicates the type of flux coating - basic, rutile etc as shown in Table 4.

Table 4 Symbol for flux coating

Symbol CoatingA acidC cellulosicR rutile

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RR thick rutileRC rutile-cellulosicRA rutile-acidRB rutile-basicB basicThe next three symbols are not compulsory and give additional information on the percentage weld metal recovery and the type of welding current on which the electrode can be operated (Table 5); the welding position (Table 6) and the maximum hydrogen content of the deposited weld metal if the electrodes are dried or baked as recommended by the manufacturer (Table 7).

Table 5 Symbol for weld metal recovery and current type

Symbol Weld metal recovery % Current type1 <= 105 AC or DC+2 <= 105 DC+ or DC-3 >105<=125 AC or DC+4 >105<=125 DC+ or DC-5 >125<=160 AC or DC+6 >125<=160 DC+ or DC-7 >160 AC or DC+8 >160 DC+ or DC-

 

Table 6 Symbols for welding position

Symbol Welding position1 All positions2 All positions except V-down3 Flat butt and fillet welds, HV fillet weld4 Flat5 V-down, flat butt, flat and HV fillet weldsTable 7 Symbol for hydrogen content in weld metal

Symbol Max Hydrogenml/100gms weld metal

H5 5H10 10H15 15A full designation may therefore read E42 2 B32H5. This describes a basic carbon manganese steel electrode; weld metal yield strength of 420N/mm2, better than 47J at -20°C, a weld metal recovery of over 105%, capable of being used on AC or DC+ current in all positions except vertical down and providing less than 5mls hydrogen in the weld metal.

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The AWS specification equivalent to EN 499 is AWS A5.1 - Carbon Steel Electrodes for Shielded Metal Arc Welding. The classification comprises five characters but in the 2004 edition of the specification there are two separate schemes. A5.1, based on the US units of tensile strength in pounds per square inch, Charpy -V values in foot-pounds and A5.1M, based on the SI system, with strength in MPa, Charpy-V values in Joules.

It is thus possible to have virtually identical electrodes with different classifications, one using US units, the other SI units. There is insufficient space within this brief article to describe fully all of the 18 types covered by the specification except perhaps for the most commonly used electrodes. For full details of the AWS scheme it is necessary to consult the specification.

To illustrate briefly how the electrodes are classified, the following gives a summary of the key features.

The first character 'E' is common to both classifications and indicates that the electrode is a flux coated manual metal arc electrode. The next two digits indicate the tensile strength. In the A5.1 designation this is either '60',indicating a UTS of 60ksi and a yield strength of 48ksi, or '70', indicating a UTS of 70ksi and a yield strength of 58ksi. In the A5.1M designation these are 43 or 49, indicating a UTS of 430MPa, yield strength of 330MPa or 490MPa UTS,400MPa yield respectively.

The last two digits give information on flux coating type, welding position, current type and polarity and Charpy-V impact strength, if required. Those electrodes suffixed XX10 or XX11 have cellulosic coatings; those suffixed XX12,XX13, XX14, XX19 or XX24 have rutile coatings and those suffixed XX15, XX16, XX18, XX28 and XX48 are basic low hydrogen. XX18, XX28 and XX48 all have iron powder additions and are therefore high recovery electrodes.

Listed below are those EN and AWS specifications that prescribe the requirements for ferrous electrodes.

BS EN 499 Non-alloyed and fine grained steel electrodesBS EN 757 High strength steelsBS EN 1599 Creep resisting steelsBS EN 1600 Stainless and heat resisting steelsAWS A5.1/A5.1M Carbon Steel Electrodes for SMAWAWS A5.4 Stainless Steel Electrodes for SMAWAWS A5.5 Low Alloy Steel Electrodes for SMAW

Welding consumables Part 4 - gas shielded consumables

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Part 1 Part 2 

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Part 3 Part 5

This article looks at the wire consumables used in the gas shielded MIG/MAG, metal cored (MC) and flux cored (FC) arc welding processes.

The MIG/MAG processes were first developed using a solid wire but some 25 years ago tubular wires began to be supplied and since then the use of these wires has grown rapidly and they now form a significant proportion of the welding wire market - cored wires are now used not only in the MIG/MAG process but also in TIG, plasma-TIG and submerged arc welding.

Solid wire for welding of alloy steels is an expensive commodity. The composition of a ferritic steel welding wire is not the same as that of the steel that it will be used to weld. The ingot from which the wire is drawn must contain all the de-oxidation and alloying elements that can be contained in the flux on an MMA electrode.

Steel is produced most economically in large tonnages whereas a consumable supplier requires only relatively small amounts and these requirements have a significant effect on the cost. In addition, it can be difficult to draw down the wire to the small diameters required for welding.

Cored wires for welding carbon and alloy steels, however, can be made from mild steel with the alloying elements added to the flux filling. This enables small amounts of wire to be economically produced matching the composition of steels where the usage is limited, eg high chromium creep resistant steels or hard facing. Non-ferrous and austenitic steel wires, aluminium, nickel based, stainless steel etc however, generally match closely the parent metal composition and obtaining ingots for drawing into wire is less of a problem.

MIG/MAG welding solid wires are provided in diameters ranging from 0.6 to 2.4mm, the most commonly used diameters being 1.2 and 1.6mm.

As mentioned above, the solid wires are generally formulated to match the composition of the alloy to be welded. Silicon, 0.5 to 0.9%, and perhaps aluminium, up to 0.15%, are added to ferritic steel wires to provide de-oxidation; carbon content is generally below 0.1%.

Alloying elements such as manganese, chromium, nickel and molybdenum are added to the ingot to provide improved mechanical properties and corrosion resistance. In addition the carbon and low alloy steel wires are often copper coated, both to reduce corrosion during storage and to improve welding current pick-up in the contact tip.

The stainless steel and non-ferrous wires are not copper coated. Poor control during the drawing operation may form laps on the wire surface that trap contaminants and give rise to porosity, as can a poor quality copper coat on ferritic steel wires.

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Porosity from drawing defects can be a particular problem with aluminium alloy wires and where high quality weld metal is required, then shaving the wire to remove defects on the wire surface is recommended.

The cored wires are small diameter tubes in which are packed fluxes and alloying elements. There are two fundamental types, one containing mostly fluxes, the other containing metal powders. There is a sub-class of the flux cored wires, the self-shielded wires, that contain gas-generating compounds that decompose in the arc to provide enough shielding gas so that additional gas shielding is not required.

In cross-section, the wires may be seamless tubes packed with the flux and extruded before being drawn into a wire. Alternatively, they may be or made by rolling a flat strip into a 'U', filling this with the flux or metal power and then folding this into a tube. The edges of the tube may be butted together or overlapped.

The seamless and closed butt wires tend to have thicker walls and therefore less fill than the overlapped wires, perhaps as little as 20% of cross sectional area compared with 50% for the overlapped wires. This enables the overlapped wires to contain more alloying elements and they are therefore often used for stainless steel and hard facing welding.

Cored wires have a number of advantages over the solid wires. The reduced current carrying cross-sectional area of the wire results in greater current density and an increase in burn-off rate with increased deposition.

The flux also produces a slag that will control weld bead shape enabling higher welding currents to be used in positional welding than can be used with MAG. A 7mm throat fillet is possible in the horizontal-vertical position, for example. The slag will also react with the weldpool and provide better properties than can be achieved with MAG. Good Charpy impact properties down to -50°C are achievable in carbon steels with the correct wire.

Disadvantages with the cored wires are:

The wire is mechanically weak and over-pressure on the wire drive rolls may crush the wire preventing it from feeding through the contact tip.

The flux cored wires produce a slag that must be removed. While solid wires often produce islands of a glassy slag that tend to lie in the finish craters this does not necessarily prevent a multi-pass weld being made without de-slagging.

This is not possible with flux cored wires, restricting their use in applications such as robotic welding to single pass welds. Metal cored wires are less of a problem in this context and are often used in fully automated, multi-pass applications.

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As with MMA electrodes, the flux in the core may be either rutile or basic, the rutile flux providing a smooth arc, easy slag removal and 'welder appeal', the basic fluxes providing better mechanical properties and cleaner radiographic quality welds.

Hydrogen control is less of a problem than with MMA electrodes. Both rutile, basic and metal cored wires all have very low hydrogen potential levels, allowing lower preheat than might otherwise be the case and enabling rutile wires to be used in applications such as the welding of high strength or thick section steels. Hydrogen pick-up on the shop floor is also less of a problem as the flux/metal powder is contained within a sealed tube, preventing moisture ingress. Seamless wires tend to be better in this respect than seamed wires.

There are a number of specifications detailing the requirements for solid and cored wires for MIG/MAG, FCA and MCA welding and these will be covered in the next article.

Welding consumables Part 5 - MIG/MAG and cored carbon steel wires

Job Knowledge

Part 1 Part 2 Part 3 Part 4

To ensure that there is a consistency in composition and properties between wires from a variety of manufacturers, specifications have been produced that enable a wire to be easily and uniquely identified by assigning the consumable a 'classification', a unique identification that is universally recognised.

The two schemes that are dealt with in this article are the EN/ISO method and the AWS scheme. There are such a large number of specifications covering the whole range of ferrous and non-ferrous filler metals, both solid wire and cored, that it will not be possible to describe all of these here. This article therefore reviews just the carbon steel specifications.

The identification of the solid wires is relatively simple, as the chemical composition is the major variable although both the EN/ISO and the AWS specifications detail the strength that may be expected from an all-weld deposit carried out using parameters given in the specification. It should be remembered, however, that most welds will contain some parent metal and that the welding parameters to be used in production may be different from those used in the test. The result is that the mechanical properties of a weld can be significantly different from those quoted by the wire supplier, hence the need to always perform a procedure qualification test when strength is important. In addition, the mechanical properties specified in the full designation include the yield strength. (In the EN/ISO specifications, the classification may indicate either yield or ultimate tensile strength).

When selecting a wire remember that the yield and ultimate tensile strengths are very close together in weld metal but can be widely separated in parent metal. A filler metal that is selected because its yield strength

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matches that of the parent metal may not, therefore, match the parent metal on ultimate tensile strength. This may cause the cross joint tensile specimens to fail during procedure qualification testing or perhaps in service.

The EN/ISO specification for non-alloyed steel solid wires is BS EN ISO 14341. This specification classifies wire electrodes in the as-welded condition and in the post weld heat-treated condition, based on classification system, strength, Charpy-V impact strength, shielding gas and composition. The classification utilises two systems based either on the yield strength (System A) or the tensile strength (System B):

System A - based on the yield strength and average impact energy of 47J of all-weld metal. System B - based on the tensile strength and the average impact energy of 27J of all-weld metal.In most cases, a given commercial product can be classified to both systems. Then either or both classification designations can be used for the product.

The symbolisation for mechanical properties is summarised in Table 1A for classification system A and Table 1B for classification system B. For classification system B, the 'X' can be either 'A' or 'P', where 'A' indicates testing in the as-welded condition and 'P' indicates testing in the post weld heat-treated condition. The symbol for chemical composition is summarised in Table 3A and 3B of BS EN ISO 14341 based on each classification system. For classification system A, the standard lists eleven compositions, too many to describe completely here. Six of the wires are carbon steel with varying amounts of deoxidants, two wires contain approximately 1% or 2.5% nickel and an additional two wires contain around 0.5% molybdenum. The designation of these wires is for example G3Si1, 'G' identifying it as a solid wire, '3' as containing some 1.5% manganese and Si1 as containing around 0.8% silicon; G3Ni1 is a wire with approximately 1.5% manganese and 1% nickel.

Table 1A Symbols for mechanical properties based on classification system A

Symbol Min Yield StrengthN/mm2

UTS

N/mm 2

Min Elongation% Symbol Charpy-V Test 47 J at Temp °C

35 355 440 to 57022 Z No requirements38 380 470 to 60020 A +2042 420 500 to 64020 0 046 460 530 to 68020 2 -2050 500 560 to 72018 3 -30        4 -40        5 -50        6 -60        7 -70        8 -80        9 -90        10 -100Table 1B Symbols for mechanical properties based on classification system B

Symbol Min Yield StrengthN/mm2

UTS

N/mm 2

Min Elongation% Symbol Charpy-V Test 27 J at Temp °C

43X 330 430 to 60020 Z No requirements49X 390 490 to 67018 Y +20

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55x 460 550 to 74017 0 057x 490 570 to 77017 2 -20        3 -30        4 -40        5 -50        6 -60        7 -70        8 -80        9 -90        10 -100A full designation could therefore be ISO 14341-A-G 46 5 M G3Si1 where the '-A' designates the classification system A, the '-G' designates solid wire electrode/or deposits, and the 'M' designates a mixed gas. An example of a System B designation could be ISO 14341-B-G 49A 6 M G3, where 'A' indicates testing in the as-welded condition.

The AWS specification AWS A5.18 covers both solid, composite stranded and cored wires comprising six carbon steel filler metals for MAG, TIG and plasma welding in both US and metric units.

The classification commences with the letters 'E' or 'ER'. 'E' designates an electrode. 'ER' indicates that the filler metal may be used either as an electrode or a rod. The next two digits designates the tensile strength in either 1000s of psi.(ksi) or N/mm2 eg ER70 (70ksi UTS) or ER48 (480N/mm2 UTS). However, note that there is only one strength level in the specification.

The next two characters identify the composition, essentially small variations in carbon, manganese and silicon contents, the wire type (solid wire (S) or metal cored or composite wire (C)) and the Charpy-V impact values.

With one exception, the solid wires are tested using 100% CO2, the cored wires with argon/CO2 or as agreed between customer and supplier, in which case there is a final letter 'C' designating CO2 or 'M', a mixed gas.

The permutations in these identifiers are too many and too complicated to be able to describe them all in sufficient detail but as an illustration, a typical designation would be ER70S-3, a 70ksi filler metal, CO2 gas shielded and with minimum Charpy-V energy of 27J at -20°C. E70C-3M identifies the wire as a solid wire 70ksi UTS metal cored filler metal, 27J at -20°C and tested with an argon/CO2 shielding gas. 

The EN/ISO specification for non-alloy steel flux and metal cored wires is BS EN ISO 17632. This covers gas shielded as well as self-shielded wires. The standard identifies electrode based on two systems in a similar way as BS EN ISO 14341, indicating the tensile properties and the impact properties of the all-weld metal obtained with a given electrode. Although the specification claims that the wires are all non-alloy, they can contain molybdenum up to 0.6% and/or nickel up to 3.85%. The classification commences with the letter 'T', identifying the consumable as a cored wire.

The classification uses the same symbols for mechanical properties as shown in Table 1A&B and a somewhat similar method to describe the composition as BS EN ISO 14341. Thus MnMo contains approximately 1.7% manganese and 0.5% molybdenum; 1.5Ni contains 1% manganese and 1.5% nickel. In addition to the symbols for properties and composition, there are symbols for electrode core composition. Table 2 summarises the symbols for electrode core type and welding position in accordance with classification system A. Classification system B uses Usability Indicators as oppose to a one-letter symbol for electrode core type, which can be found in Table 5B of BS EN ISO 17632.

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Table 2 Symbols for electrode core type and position based on classification system A

Flux Core Welding PositionSymbol Flux Core Type Shielding Gas Symbol Welding positionR Rutile, slow freezing slag Required 1 AllP Rutile, fast freezing slag Required 2 All except V-downB Basic Required 3 Flat butt, flat and HV filletM Metal powder Required 4 Flat butt and filletV Rutile or basic/fluoride Not required 5 V-down and (3)W Basic/fluoride, slow freezing slag Not required    Y Basic/fluoride, fast freezing slag Not required    Z Other types      In addition, there are symbols for gas type. These are 'M' for mixed gases, 'C' for 100% CO2 and 'N' for self-shielded wires and 'H' for hydrogen controlled wires. A full designation may therefore be ISO 17632-A -T46 3 1Ni B M 1 H5 in accordance with classification system A. For classification system B, an example may be ISO 17632-B -T55 4 T5-1MA-N2-UH5, where 'T5' is the usability designator, 'A' indicates test in the as-welded condition, 'N2' is the chemical composition symbol, and 'U' is an optional designator.

The American Welding Society classification scheme for carbon steel flux cored wires is detailed in the specification AWS A5.36. This also contains information from A5.18, but does not officially supercede it. The full designation is ten characters in length beginning 'E' for an electrode then designators for strength, welding position, cored wire, usability, shielding gas, toughness, heat input limits and diffusible hydrogen, the last four designators being optional.

There are two strength levels - E7 (70ksi UTS) and E6 (60ksi UTS) followed by a designator for welding position,'0' for flat and horizontal and '1' for all positions, including vertical-up and vertical-down.

The next symbol 'T' identifies the wire as being flux cored and this is followed by either a number between 1 and 14 or the letter 'G' that identifies the usability. This number refers to the recommended polarity, requirements for external shielding, and whether the wire can be used to deposit single or multi-pass welds. 'G' means that the operating characteristics are not specified. The sixth letter identifies the shielding gas used for the classification, 'C' being 100% CO2, 'M' for argon/CO2, no letter indicating a self-shielded wire.

The non-compulsory part of the designation may include the letter 'J', confirming that the all-weld metal test can give Charpy-V values of 27J at -40°C; the next designator may be either 'D' or 'Q'. These indicate that the weld metal will achieve supplementary mechanical properties at various heat inputs and cooling rates. The final two designators identify the hydrogen potential of the wire.

A full AWS A5.36 designation could therefore be E71T-2M-JQH5. This identifies the wire as a cored, all positional wire to be used with argon/CO2 shielding gas on electrode positive polarity. The weld metal should achieve 70ksi tensile strength, 27J at -40°C, 58 to 80ksi yield strength at high heat input, a maximum 90ksi at low heat input, and a diffusible hydrogen content of less than 5ml of H2/100g of deposited weld metal.

Submerged arc welding consumables - Part 1

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268

The submerged arc process is somewhat unusual in that the welding consumables, unlike the other fluxed processes of MMA or FCAW, comprise two components, the wire and the flux, that may be supplied separately.

Since both the wire and the flux will have an effect on the weld metal composition, and hence on the mechanical properties, the welding engineer is faced with choosing the appropriate wire/flux combination for the application. This article discusses some of the characteristics of wires and fluxes. The next article will review the specifications.

The welding wire is generally of a composition that matches that of the parent metal and wires are available for the welding of carbon and low and high alloy steels, stainless steels, nickel and copper/nickel alloys. In addition, submerged arc welding may be used for surfacing with corrosion or wear resistant coatings using both wires and flat strips. The wires may be solid or metal cored. Strips may be rolled or sintered.

Welding wires vary from 1.2mm ('thin' wire or twin wire submerged arc) to 6.4mm in diameter and are capable of carrying welding currents ranging from 150 to 1600amps. The wires for ferritic steels are generally copper coated to increase contact tip life, improve electrical conductivity and extend the shelf life. Stainless steel and nickel alloy wires are bright drawn and uncoated. The wire is supplied on reels weighing 10 to 50kg and can also be obtained in large pay-off packs weighing up to 500kg. The strip used for surfacing is supplied in 15 to 240mm widths but the thickness is a standard 0.5mm. As with the wire, strip is available in a range of coil weights.

Whilst the wire is relatively simple and is designed to match the parent metal composition and/or mechanical properties, the flux is far more complex. The functions of the flux are:

to assist arc striking and stability to form a slag that will protect and shape the weld bead to form a gas shield to protect the molten filler metal being projected across the arc gap to react with the weld pool to provide clean high quality weld metal with the desired properties to deoxidise the weld pool provide deoxidants in some circumstances, to provide additional alloying elements into the weld poolFluxes may be categorised in two ways: by the method of manufacture (fused or agglomerated) or by its activity (neutral, active or alloying). Within these broad groupings the fluxes may be classified further by their constituents, silica, manganese oxide, calcium fluoride etc.

Perhaps the most convenient method of classifying, however, is by reference to the 'basicity index' (BI) of the flux. The index is calculated by dividing the sum of the percentages of the basic constituents by the sum

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of the acid constituents. Calcium, magnesium, sodium, potassium and manganese oxides, calcium carbonate and calcium fluoride are the basic constituents of a flux; silica and alumina the acid constituents. Acid fluxes have a basicity index of 0.5 to 0.8; neutral fluxes 0.8 to 1.2; basic fluxes 1.2 to 2.5 and highly basic fluxes 2.5 to 4.0. The basicity of a flux has a major effect on the weld metal properties, most importantly the notch toughness. As a general rule the higher the basicity the higher the notch toughness.

Neutral fluxes are designed to have little or no effect on the chemical analysis of the weld metal and therefore on the mechanical properties. They contain low silica, calcium silicate and alumina and do not add significant amounts of silicon and manganese to the weld.

The acid fluxes contain substantial amounts of silica, silicates in the form of calcium and/or manganese silicate and manganese oxide. These fluxes react with the weld pool and will raise both silicon and manganese content of the weld together with a high oxygen content. The result of this is that the toughness of the weld is poor but the fluxes will tolerate rusty surfaces, will detach easily and give a good weld appearance. They are especially useful for single pass high speed welding such as fillet welding of web to flange girder joints.

The basic fluxes fill much the same role in submerged arc welding as basic coatings do in manual metal arc welding. They have a low silica content and are composed of varying amounts of calcium carbonate and/or fluoride, alumina, calcium, manganese and magnesium oxides and rutile.

This combination of compounds gives a clean, low sulphur, low oxygen weld metal with good to excellent notch toughness. As a general rule, the higher the basicity, the higher the toughness. The transfer of silicon and manganese into the weld metal is also limited. Such fluxes are preferred for the welding of high quality structural steels, pressure vessels, pipework and offshore structures where either good high or low temperature properties are required.

The fused fluxes are acid, neutral or slightly basic and are manufactured by mixing the constituents together, melting them in an electric furnace and crushing the solidified slag that is produced to give a flux with a glassy appearance.

These fluxes are homogeneous, resistant to moisture pick-up and mechanically strong so that they do not break down but maintain the required particle size. The high temperatures required by the melting operation mean that some constituents, particularly the de-oxidants present in the highly basic fluxes, decompose and are lost. This limits the range of applications of these fluxes to general structural work where sub-zero service temperatures will not been countered.

The agglomerated fluxes may be neutral, basic or highly basic. They are made from a wet mix that is corned, dried and baked to achieve a low moisture content. This low temperature process means that strong deoxidants and ferro-alloys can be incorporated without being lost. The binders used in the corning process, however, are hygroscopic so that moisture pick-up can be a problem on the shop floor. Baking of the flux prior to use may be necessary and the flux should be stored on the welding equipment in heated hoppers. The flux may also suffer mechanical damage during recirculation, breaking down to form a dust. Although a small particle size is capable of carrying a higher current, too many fines in the flux will give rise to gas being trapped between the slag and the weld pool. This will result in unsightly gas flats or pockmarking on the weld surface. To avoid this, the recirculating system should be equipped with filters to remove both large particles of detached slag and the fine dust.

Fluxes are supplied in bags, generally plastic, weighing from 25 to 40kg and in plastic drums of up to 250kg. Recently some suppliers have been packing the flux in hermetically sealed bags, aka vacuum packed electrodes. This method is useful in that the flux can be used straight from the bag with guaranteed low hydrogen levels and without the need to bake prior to use.

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Submerged arc welding consumables. Part 2 - specifications

Job Knowledge 

Part 1 Part 3

Of all the arc welding processes, only submerged arc welding uses two completely separate components, both of which may have a major effect on the mechanical properties of the weld deposit. This makes the specifying of consumables somewhat complicated. It will not be possible therefore to cover all the alloy types in this brief article which will cover the carbon, carbon-manganese and low alloy structural steels only.

BS EN ISO 14171:2010 is the specification for Welding consumables: Solid wire electrodes, tubular cored electrodes and electrode/flux combinations for submerged arc welding of non alloy and fine grain steels. This standard replaces BS EN 756:2004.

The specification covers the classification of the wire chemical composition and the wire/flux combination. It also specifies the mechanical properties of all weld metal deposits in the as-welded condition.

This standard is a combined specification providing for classification utilizing a system based upon the yield strength and the average impact energy for weld metal of 47 J, or utilizing a system based upon the tensile strength and the average impact energy for weld metal of 27 J.

The classification is composed of:

A reference to the standard 'ISO 14171' A symbol 'A' if the classification is based on yield strength and average impact energy is 47J or 'B' if

the classification is based on tensile strength average impact energy is 27J.And of five parts, plus a sixth supplementary part:

Part 1. A symbol indicating the process - in the case of submerged arc welding this is 'S'.Part 2. Two digits indicating either the tensile properties of a multi-run deposit or the tensile properties of the parent metal to be welded using a two run technique - see Tables 1 and 2.

Table 1A. Symbols for tensile properties - multi-run technique (classification based on yield strength and average impact energy 47J)

Multi-run Tensile PropertiesSymbol Min. Yield N/mm 2 Min. UTS N/mm 2 Min. Elongation %

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35 355 440 - 570 2238 380 470 - 600 2042 420 500 - 640 2046 460 530 - 680 2050 500 560 - 720 18

Table 1B. Symbols for tensile properties - multi-run technique (classification based on tensile strength and average impact energy 27J)

Multi-run Tensile PropertiesSymbol Min. Yield N/mm 2 Min. UTS N/mm 2 Min. Elongation %

43X 330 430 - 600 2049X 390 490 - 670 1855X 460 550 - 740 1757X 490 570 - 770 17

Note: 'X' is 'A' or 'P', where 'A' indicates testing in the as-welded condition and 'P' indicates testing in the post-weld heat-treated condition.

Table 2A. Symbols for tensile properties - two-run technique (classification based on yield strength and average impact energy 47J)

Two-Run Tensile PropertiesSymbol Min. Yield Parent Metal N/mm 2 Min. Tensile Strength of Welded Joint N/mm 2

2T 275 3703T 355 4704T 420 5205T 500 600

Table 2B. Symbols for tensile properties - two-run technique (classification based on tensile strength and average impact energy 27J)

Symbol Min. Tensile Strength of Welded Joint N/mm 2

43S 43049S 49055S 55057S 570

Note that the two-run technique has two tensile results specified; one for the minimum yield strength of the parent metal, one for the tensile strength of the welded joint.

Part 3. Table 3 gives the temperature at which the average Charpy-V impact value of 47J or 27J may be achieved.

Table 3. Symbol for Charpy-V impact properties

Symbol Temp. for Min Impact Energy 47J or 27J at °CZ No requirementsA +200 02 -20

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3 -304 -405 -506 -607 -708 -809 -9010 -100

Part 4. The symbol for welding flux type shall be in accordance with ISO 14174.

Flux type symbol

Flux Type Symbolmanganese-silicate MScalcium-silicate CSzirconium-silicate ZSrutile-silicate RSaluminate-rutile ARaluminate-basic ABaluminate-silicate ASaluminate-fluoride basic AFfluoride-basic FBany other type ZPart 5. Tables 4 and 5 in ISO 14171 contain a listing of the chemical composition of 22 wires and are too lengthy to include in full in this article. The wires all contain a maximum carbon content of 0.15% and range from plain carbon, through C-Mn, C-Mo, Mn-Mo to Ni and Ni-Mo. All are prefixed 'S' followed by a number from 1 to 4 denoting from 0.5% Mn (1) to 2% Mn (4). The addition of nickel and/or molybdenum is denoted by the chemical symbol of the alloy addition being included. Thus an S3 wire contains 1.5% Mn, an S2Ni1Mo 1%.

Part 6. (optional) The standards also provides symbols symbols given an optional symbol indicating the diffusible hydrogen content of the weld metal obtained in accordance with ISO 3690 (see Table 6 in the standard).

Examples of designations:

The designation for an electrode/flux combination for submerged arc welding for multi-run technique depositing a weld metal with a minimum yield strength of 460 MPa (46) and a minimum average impact energy of 47 J at -30°C (3) produced with an aluminate-basic flux (AB) and a wire S2 would be:

ISO 14171-A-S 46 3 AB S2

In addition to BS EN ISO 14171 which specifies the mechanical properties expected from a particular flux/wire combination, there is an additional specification, BS EN ISO 14174:2012, that specifies the fluxes in greater detail, including the application for which a flux may be used. The specification uses a total of seven symbols, four being compulsory and three optional. The first symbol,'S', identifies the flux as being intended for submerged arc welding and the second the method of manufacture. This may be 'F', a fused flux; 'A', an agglomerated flux and 'M', a mixture of fused and agglomerated. The third part gives an

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indication of the chemical constituents and uses the same notation as in Table 4 above. In addition BS EN 760 gives a range of percentages for each of the constituents in Table 1.

The fourth part gives a symbol for the application(s), Class 1 being intended for the welding of carbon and low alloy steels, including high strength structural and creep resistant steels. There is no alloying from this class of flux. Class 2 fluxes are for the welding of, and the surfacing with, stainless and heat resisting steels and nickel alloys. Class 3 is for use with hard surfacing weld metals, the flux providing such elements as carbon, chromium and molybdenum to the weld deposit.

The remaining three symbols are not compulsory and comprise, firstly, a number or chemical element symbol that defines what is termed in the specification as the 'metallurgical behaviour' of the three classes of flux mentioned above. Two digits then specify the pick-up or loss of silicon and manganese (in this order) to be expected when welding carbon or low alloy steels using flux Class 1, as shown in Table 5 below. Flux Classes 2 and 3 may be characterised by the use of a chemical symbol to identify the alloying element being added via the flux, eg Cr, if the flux is chromium compensating.

The current type is indicated by the addition of DC or AC to the symbols and finally an 'H', followed by a number, gives the weld metal hydrogen level expected from a correctly dried or baked flux eg H5.

A designation for a flux supplied in accordance with BS EN 760 may therefore be S A AF 1 55 DC H5 for an agglomerated alumina-calcium fluoride basic flux intended for the welding of carbon or low alloy steels, no pick-up or loss of silicon or manganese, used with DC welding current and with a hydrogen content of less than 5mls/100gms weld metal.

It must be remembered that the properties given by these designations are obtained from as welded, all weld metal specimens deposited using standard welding parameters of current, voltage and travel speed.

The properties achieved in a production weld may be entirely different due to the effects of dilution from the parent metal, higher or lower heat input, different wire diameters, preheat and interpass temperatures and post weld heat treatment. It is essential, therefore, that the suitability of a flux/wire combination is confirmed by procedure qualification testing.

Note also that flux/wire combinations supplied to the same specification designation by different manufacturers may not necessarily provide similar mechanical properties or weld cleanliness.

Submerged arc welding consumables. Part 3 - AWS specifications

Job Knowledge

Part 1 Part 2

As with the BS EN specifications for submerged arc welding consumables, the American Welding Society (AWS) system also uses a dual flux type/wire composition designation to identify the flux/wire combination that will provide the required properties.

The AWS system is somewhat simpler than the BS EN method, particularly if the full flux descriptor is used as specified in BS EN 760 (see Connect article No. 88). There are, however, only two specifications that deal with both wire composition and the flux but an additional two specifications that cover bare wires for stainless steels and the nickel based alloys. These are ANSI/AWS A5.17 - Carbon Steel Electrodes and Fluxes and ANSI/AWS A5.23 Low Alloy Steel Electrodes and Fluxes. The bare wire specifications are ANSI/AWS A5.9 Bare Stainless Steel Welding Electrodes and Rods and ANSI/AWS A5.Nickel and Nickel Alloy Bare Welding Electrodes and Rods.

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In AWS A.5.17 and AWS A5.23 the first part of the designation describes the flux type and may comprise up to six digits depending upon whether the flux is supplied with the tensile strength expressed in increments of 10 megapascals (two numbers where 43 represents 430MPa) or in pounds per square inch (1 digit ie 6 represents 60,000psi).

The first digit, the letter 'F', identifies the consumable as a submerged arc welding flux, the next letter 'S' is only included if the flux is made from or includes crushed slag. Omission of this letter 'S' indicates that the flux is unused and contains no crushed used flux introduced either by the flux manufacturer or the welding fabricator.

The next one or two digits specify the minimum tensile strength as explained above and this is followed by 'A' or 'P' for whether the test results were obtained in the as-welded, the A condition or post-weld heat treated, the P condition. The last digit identifies the minimum temperature at which a Charpy-V impact value of 27J can be achieved as in Table 1 below.

Table 1 Impact Test Requirements

DigitTest Temperature Impact value

Joules°C °FZ no impact requirements 270 -18 0 272 -29 -20 274 -40 -40 275 -46 -50 276 -51 -60 278 -62 -80 27

In AWS A5.17 there is a total of eleven wires, split into three groups of low, medium and high manganese. The first digit, 'E', identifies the consumable as a bare wire electrode. If supplemented by 'C' the wire is a composite (cored) electrode. The composition of the solid wire is obtained from an analysis of the wire. However, since the composition of a cored wire may be different from that of its weld deposit the composition must be determined from a low dilution weld deposit made using a specific, named flux.

The next letter, 'L', 'M' or 'H' indicates a low (0.6% max), medium (1.4% max) or high (2.2% max) manganese content. This is followed by one or two digits that give the nominal carbon content. An optional letter 'K' indicates a silicon killed steel. There are a final two or three optional digits identifying the diffusible hydrogen in ml/100gms weld metal, H16, H8 or H4.

A full designation for a carbon steel flux/wire combination could therefore be F6P5-EM12K-H8. This identifies this as being a solid wire with a nominal 0.12% carbon, 1% manganese and 0.1 to 0.35% silicon capable of achieving an ultimate tensile strength of 60 k.p.i. (415MPa), a Charpy-V impact strength of 27J at -50°F (-46°C) in the post weld heat treated condition.

The classification in AWS A5.23 is, of necessity, rather more complicated as this specification covers a wide range of low alloy steels, a total of thirty one solid wires and twenty nine composite wire weld metal compositions. Within the confines of this brief article it will not be possible to cover in full the entire classification of the wires.

The flux designation is almost identical to that of AWS A5.17, except that a four, five or six digit identifier may be used. Why this additional sixth digit? Because some of the electrodes in the specification are

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capable of providing tensile strengths above 100,000 psi - in these cases the designation may be, for example, F11, identifying the flux as providing 110 ksi (760MPa) minimum tensile strength.

The classification of the wire comprises two parts - the first that of the wire, solid wires being prefixed 'E' and composite wires 'EC', the second part specifies the composition of the weld deposit. In Table 1 of the specification only the solid wires are listed. The wire classification commences with 'E' to identify a bare wire, the next letter places the wire in a 'family' of wires. 'L' or 'M' identifies the wires as being alloyed with copper, 0.35% max; 'A'as containing molybdenum, 0.65% max; 'B' as the creep resisting steels containing chromium and molybdenum; 'Ni' for those wires containing nickel. 'F comprises the Ni-Mo or Cr-Ni-Mo wires; 'M' triple de-oxidised Ni-Mo wires; 'W' aNi-Cu wire and 'G' not specified.

This use of wires to this latter 'G' designation may lead to problems as quite large changes can be made to the composition to achieve the required mechanical properties - a good example of this is where the NACE requirements for sour service of 248BHN or 1% nickel maximum are required. To achieve the required tensile or impact strength the consumable manufacturer may increase the carbon or nickel contents above those used in the procedure qualification test and still supply to the same designation.

Table 2 in AWS A5.23 classifies both solid wire and composite wire/flux combinations by means of weld metal compositions but still using the identifying letters as for the solid wires described above. The prefix 'E' is, however, omitted thus a carbon/molybdenum deposit may be classified, for example, as A3, a Cr-Mo deposit as B4, Ni-Mo as F5 etc.

Thus a full designation for a flux/wire combination for an as welded 1% Ni/0.25% Mo weld deposit with an ultimate tensile strength of 80ksi and an impact strength of 27J at -60°F (-51°C) may therefore be F8A6-ENi1-Ni1 and for a similar deposit using a cored wire in the PWHT'd condition F8P6-ECNi1-Ni1.

As mentioned in earlier articles on the topic of consumable specifications, it must be remembered that the mechanical properties and compositions are determined from test pieces taken from absolutely minimal dilution welds made on specified parent plates with a standard set of welding parameters - heat input, preheat, interpass temperature, post weld heat treatment temperature and time. They may therefore NOT reflect the results obtained in a production weld and the designation cannot be relied upon to guarantee the properties required by the application.

Where these properties are important it is therefore essential that mechanical testing, chemical analysis etc are determined from test specimens made using parent materials and parameters representative of production welding.

Heat treatment of welded joints

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Job Knowledge

Heat treatment is an operation that is both time consuming and costly. It can affect the strength and toughness of a welded joint, its corrosion resistance and the level of residual stress but is also a mandatory operation specified in many application codes and standards. In addition it is an essential variable in welding procedure qualification specifications.

Before discussing the range of heat treatments that a metal may be subjected to, there is a need to clearly define what is meant by the various terms used to describe the range of heat treatments that may be applied to a welded joint. Such terms are often used incorrectly, particularly by non-specialists; for a metallurgist they have very precise meanings.

Fig. 1 Heat treatment of welded joints

Solution treatment

Carried out at a high temperature and designed to take into a solution elements and compounds which are then retained in solution by cooling rapidly from the solution treatment temperature. This may be done to reduce the strength of the joint or to improve its corrosion resistance. With certain alloys it may be followed by a lower temperature heat treatment to reform the precipitates in a controlled manner  (age or precipitation hardening).

Annealing

This consists of heating a metal to a high temperature, where recrystallisation and/or a phase transformation take place, and then cooling slowly, often in the heat treatment furnace. This is often carried out to soften the metal after it has been hardened, for example by cold working; a full anneal giving the very softest of microstructures. It also results in a reduction in both the yield and the tensile strength and, in the case of ferritic steels, usually a reduction in toughness.

Normalising

This is a heat treatment that is carried out only on ferritic steels. It comprises heating the steel to some 30-50°C above the upper transformation temperature (for a 0.20% carbon steel this would be around 910°C) and cooling in still air. This results in a reduction in grain size and improvements in both strength and toughness.

Quenching

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This comprises a rapid cool from a high temperature. A ferritic steel would be heated to above the upper transformation temperature and quenched in water, oil or air blast to produce a very high strength, fine grained martensite. Steels are never used in the quenched condition, they are always tempered following the quenching operation.

Tempering

A heat treatment carried out on ferritic steels at a relatively low temperature, below the lower transformation temperature; in a conventional structural carbon steel this would be in the region of 600-650°C. It reduces hardness, lowers the tensile strength and improves ductility and toughness. Most normalised steels are tempered before welding, all quenched steels are used in the quenched and tempered condition.

Ageing or Precipitation hardening

A low temperature heat treatment designed to produce the correct size and distribution of precipitates, thereby increasing the yield and tensile strength. It is generally preceded by a solution heat treatment. For steel, the temperature may be somewhere between 450-740 degree C, an aluminium alloy would be aged at between 100-200°C. Longer times and/or higher temperatures result in an increase in size of the precipitate and a reduction in both hardness and strength.

Stress relief

As the name suggests, this is a heat treatment designed to reduce the residual stresses produced by weld shrinkage. It relies upon the fact that, as the temperature of the metal is raised, the yield strength decreases, allowing the residual stresses to be redistributed by creep of the weld and parent metal. Cooling from the stress relief temperature is controlled in order that no harmful thermal gradients can occur.

Post heat

A low temperature heat treatment carried out immediately on completion of welding by increasing the preheat by some 100°C and maintaining this temperature for 3 or 4 hours. This assists the diffusion of any hydrogen in the weld or heat affected zones out of the joint and reduces the risk of hydrogen induced cold cracking. It is used only on ferritic steels, where hydrogen cold cracking is a major concern i.e. very crack sensitive steels, very thick joints etc.

Post Weld Heat Treatment (PWHT)

So what does the term 'post weld heat treatment' mean? To some engineers it is a rather vague term that is used to describe any heat treatment that is carried out when welding is complete. To others however, particularly those working in accordance with the pressure vessel codes such as BS PD 5500, EN 13445 or ASME VIII, it has a very precise meaning. When an engineer talks of post weld heat treatment, annealing, tempering or stress relief it is therefore advisable. 

Heat treatment following welding may be carried out for one or more of three fundamental reasons:

to achieve dimensional stability in order to maintain tolerances during machining operations or during shake-down in service

to produce specific metallurgical structures in order to achieve the required mechanical properties to reduce the risk of in-service problems such as stress corrosion or brittle fracture by reducing the

residual stress in the welded component The range of heat treatments to achieve one or more of these three objectives in the range of ferrous and non-ferrous metals and alloys that may be welded is obviously far too extensive to cover in great detail

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within these brief Job Knowledge articles. The emphasis in the following section will be on the PWHT of carbon and low alloy steels as required by the application standards although brief mention will be made of other forms of heat treatment that the welding engineer may encounter in the ferrous alloys. There are two basic mechanisms that are involved, firstly stress relief and secondly microstructural modifications or tempering.

Stress Relief

Why is it necessary to perform stress relief? It is an expensive operation requiring part or all of the welded item to be heated to a high temperature and it may cause undesirable metallurgical changes in some alloys. As mentioned above there may be one or more reasons. The high residual stresses locked into a welded joint may cause deformation outside acceptable dimensions to occur when the item is machined or when it enters service. High residual stresses in carbon and low alloy steels can increase the risk of brittle fracture by providing a driving force for crack propagation. Residual stresses will cause stress corrosion cracking to occur in the correct environment eg carbon and low alloy steels in caustic service or stainless steel exposed to chlorides.

What causes these high residual stresses? Welding involves the deposition of molten metal between two essentially cold parent metal faces. As the joint cools the weld metal contracts but is restrained by the cold metal on either side; the residual stress in the joint therefore increases as the temperature falls. When the stress has reached a sufficiently high value (the yield point or proof strength at that temperature) the metal plastically deforms by means of a creep mechanism so that the stress in the joint matches the yield strength. As the temperature continues to fall the yield strength increases, impeding deformation, so that at ambient temperature the residual stress is often equal to the proof strength (Fig 1).                

To reduce this high level of residual stress, the component is reheated to a sufficiently high temperature. As the temperature is increased the proof strength falls, allowing deformation to occur and residual stress to decrease until an acceptable level is reached. The component would be held at this temperature (soaked) for a period of time until a stable condition is reached and then cooled back to room temperature. The residual stress remaining in the joint is equal to the proof strength at the soak temperature.

Figure 1 shows that residual stress in a carbon manganese steel falls reasonably steadily from ambient to around 600 degree C but that the high strength creep resistant steels need to be above 400 degree C before the residual stress begins to fall. Stainless steel is hardly affected until the temperature exceeds 500 degree C. There is therefore a range of soak temperatures for the various alloys to achieve an acceptable reduction in residual stress without adversely affecting the mechanical properties of the joint. In carbon manganese steels this temperature will be between 550-620 degree C, in creep resistant steels somewhere between 650-750 degree C and for stainless steels between 800-850 degree C.

The next article will cover tempering of ferritic steels and will be followed by further information on other alloys and methods of applying and controlling heat treatment activities.  

Heat treatment of welded joints - Part 2

Job Knowledge

Part 1 of this series of articles gave definitions of some of the heat treatments that may be applied to a welded joint and dealt with the operation of stress relieving a ferritic steel assembly. The temperature range within which stress relief takes place will also cause tempering of those regions in the HAZ’s where hard structures may have formed.

Tempering

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Tempering is a heat treatment that is only relevant to steels and is carried out to soften any hard micro-structures that may have formed during previous heat treatments, improving ductility and toughness. Tempering also enables precipitates to form and for the size of these to be controlled to provide the required mechanical properties. This is particularly important for the creep resistant chromium-molybdenum steels. Tempering comprises heating the steel to a temperature below the lower critical temperature; this temperature being affected by any alloying elements that have been added to the steel so that for a carbon-manganese steel, the temperature is around 650°C, for a 2¼CrMo steel, 760°C . Quenched steels are always tempered. Normalised steels are also usually supplied in the tempered condition although occasionally low carbon carbon-manganese steel may be welded in the normalised condition only, the tempering being achieved during PWHT. Annealed steels are not supplied in the tempered condition.

Tempering of tool steels may be performed at temperatures as low as 150 degrees C, but with the constructional steels that are the concern of the welding engineer the tempering temperature is generally somewhere between 550- 760°C, depending on the composition of the steel.

Post Weld Heat Treatment (PWHT)

As mentioned in Part 1, PWHT is a specific term that encompasses both stress relief and tempering and is not to be confused with heat treatments after welding. Such treatments may comprise ageing of aluminium alloys, solution treatment of austenitic stainless steel, hydrogen release etc. PWHT is a mandatory requirement in many codes and specifications when certain criteria are met. It reduces the risk of brittle fracture by reducing the residual stress and improving toughness and reduces the risk of stress corrosion cracking. It has, however, little beneficial effect on fatigue performance unless the stresses are mostly compressive.

It is an essential variable in all of the welding procedure qualification specifications such as ISO 15614 Part 1 and ASME IX. Addition or deletion of PWHT or heat treatment outside the qualified time and/or temperature ranges require a requalification of the welding procedures. PWHT temperatures for welds made in accordance with the requirements of EN 13445, ASME VIII and BS PD 5500 are given below in Table 1.

Table 1: PWHT Temperatures from Pressure Vessel Specifications 

 Steel Grade  BS EN 13445  ASME VIII  BS PD 5500

 

  Temp range°C

  Normal holding temp °C

  Temp range °C

 C Steel  550-600  593  580-620 C 1/2 Mo  550-620  593  630-670 1Cr 1/2 Mo  630-680  593  630-700 2 1/4 Cr/Mo  670-720  677  630-750 5CrMo  700-750  677  710-750 3 1/2 Ni  530-580  593  580-620Note from Table 1 that ASME VIII specifies a minimum holding temperature and not a temperature range as in the BS and EN specifications.

As mentioned above, PWHT is a mandatory requirement when certain criteria are met, the main one being the thickness. BS EN 13445 and BSPD 5500 require that joints over 35mm thick are PWHT’d, ASME VII above 19mm. If, however, the vessel is to enter service where stress corrosion is a possibility, PWHT is mandatory, irrespective of thickness. The soak time is also dependant on thickness. As a very general rule this is one hour per 25mm of thickness; for accuracy, reference must be made to the relevant specification.

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These different requirements within the specifications mean that great care needs to be taken if a procedure qualification test is to be carried out that is intended to comply with more than one specification. A further important point is that the PWHT temperature should not be above that of the original tempering temperature as there is a risk of reducing the strength below the specified minimum for the steel. It is possible to PWHT above the tempering temperature only if mechanical testing is carried out to show that the steel has adequate mechanical properties. The testing should, obviously, be on the actual material in the new heat treatment condition.

Maximum and minimum heating and cooling rates above 350-400°C are also specified in the application codes. Too fast a heating or cooling rate can result in unacceptable distortion due to unequal heating or cooling and, in very highly restrained components, may cause stress cracks to form during heating.

Application of PWHT

The method of PWHT depends on a number of factors; what equipment is available, what is the size and configuration of the component, what soaking temperature needs to be achieved, can the equipment provide uniform heating at the required heating rate? The best method is by using a furnace. This could be a permanent fixed furnace or a temporary furnace erected around the component, this latter being particularly useful for large unwieldy structures or to PWHT a large component on site. Permanent furnaces may be bogie loaded with a wheeled furnace bed on to which the component is placed or a top hat furnace that uses a fixed hearth and a removable cover. Typically, a furnace capable of heat treating a 150tonne pressure vessel would have dimensions of around 20m long, a door 5x5m and would consume around 900cu/metres of gas per hour.

Furnaces can be heated using electricity, either resistance or induction heating, natural gas or oil. If using fossil fuels care should be taken to ensure that the fuel does not contain elements such as sulphur that may cause cracking problems with some alloys, particularly if these are austenitic steels or are nickel based – corrosion resistant cladding for example. Whichever fuel is used the furnace atmosphere should be closely controlled such that there is not excessive oxidation and scaling or carburisation due to unburnt carbon in the furnace atmosphere. If the furnace is gas or oil fired the flame must not be allowed to touch the component or the temperature monitoring thermocouples; this will result in either local overheating or a failure to reach PWHT temperature.

Monitoring the temperature of the component during PWHT is essential. Most modern furnaces use zone control with thermocouples measuring and controlling the temperature of regions within the furnace, control being exercised automatically via computer software. Zone control is particularly useful to control the heating rates when PWHT’ing a component with different thicknesses of steel. It is not, however, recommended to use monitoring of the furnace temperature as proving the correct temperatures have been achieved in the component. Thermocouples are therefore generally attached to the surface of the component at specified intervals and it is these that are used to control the heating and cooling rates and the soak temperature automatically so that a uniform temperature is reached. There are no hard and fast rules concerning the number and disposition of thermocouples, each item needs to be separately assessed.

As mentioned earlier, the yield strength reduces as the temperature rises and the component may be unable to support its own weight at the PWHT temperature. Excessive distortion is therefore a real possibility.  It is essential that the component is adequately supported during heat treatment and trestles shaped to fit the component should be placed at regular intervals. The spacing of these will depend on the shape, diameter and thickness of the item. Internal supports may be required inside a cylinder such as a pressure vessel; if so, the supports should be of a similar material so that the coefficients of thermal expansion are matched.

Whilst heat treating a pressure vessel in one operation in a furnace large enough to accommodate the entire vessel is the preferred method this is not always possible. In this case the pressure vessel application codes

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permit a completed vessel to be heat treated in sections in the furnace. It is necessary to overlap the heated regions – the width of the overlap is generally related to the vessel thickness. BS EN 13445 for instance specifies an overlap of 5√Re where R = inside diameter and e = thickness; ASME VIII specifies an overlap of 1.5 metres. It should be remembered that if this is done there will be a region in the vessel (which may contain welds) that will have experienced two cycles of PWHT and this needs to be taken into account in welding procedure qualification testing. There is also an area of concern, this being the region between the heated area within the furnace and the cold section outside the furnace. The temperature gradient must be controlled by adequately lagging the vessel with thermally insulating blankets and the requirements are given in the application codes.

It is, of course, possible to assemble and PWHT a vessel in sections and then to carry out a local PWHT on the final closure seam. Local PWHT will be discussed in the next part of this series on heat treatment.

The next article will cover further information on other alloys and methods of applying and controlling heat treatment activities. 

Heat Treatment Part 3

Job Knowledge

When it is not possible to place the entire component in a furnace for heat treatment (because of the size of the fabrication, circumferential welds in a pipework system or when installing equipment on site, for example), then a local PWHT may be the only option. Local PWHT needs careful planning to ensure that heating and cooling rates are controlled and that an even and correct temperature is achieved. Uneven and/or rapid heating can give rise to harmful temperature gradients producing thermally induced stresses that exceed the yield stress. This may result in the development of new residual stresses when the component is cooled. 

Local PWHT may be carried out using high velocity gas burners, infra red burners, induction heating and high or low resistance heating elements. Electrical equipment is more easily installed and controlled than heating using natural gas or propane, particularly on site. High voltage resistance heating is rarely used on site due to the need for the radiant heaters to be positioned a set distance from the surface and, more significantly perhaps, the health and safety risks involved with the use of high voltage current. Low voltage electrical resistance heating and induction heating are the two most commonly used methods.

High velocity gas burners are more advantageous when large areas need to be heat treated, particularly if, for example, firing can take place within a pressure vessel which then becomes its own furnace. For local PWHT of vessel circumferential seams internal insulating barriers can be used to localise the heat source. Motorised valves and micro-processor control of the combustion conditions enabled precise management of the heating cycle to be achieved.

Low voltage electrical resistance heating uses flexible ceramic heating elements, colloquially known as corsets, an appropriate number being assembled to cover the area to be heat treated. Induction heating uses insulated cables that can be wrapped around the joint or shaped to fit the area to be heated or specially designed fitting for repetitive PWHT operations as illustrated in Fig 1. To perform the PWHT, temperature control thermocouples are firstly attached, often by capacitor discharge welding, the elements placed in position and the area then lagged with thermal insulating blankets to reduce heat loss and to maintain an acceptable temperature gradient.

There are no standard terms used to describe the various regions within the locally PWHT'd area. In this article the terms 'soak band', 'heated band', 'gradient control band', 'temperature gradient', which may be axial and through thickness, and 'control zone' as suggested by the ASME will be used (see Fig 2).

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The soak band is the area that is heated to, within the specified PWHT temperature and time range. It comprises the weld, the two HAZs and part of the surrounding parent metal. The heated band is the area covered by the heating elements, the temperature at the edge of the heated band generally being required to be at least half that of the soak temperature. The temperature gradient control band is the region where thermal insulation, perhaps supplemented by additional heating elements, is applied to ensure that an acceptable axial temperature gradient is achieved from PWHT temperature to ambient. A control zone is the region where a number of heating elements are grouped together and controlled by a single thermocouple, enabling different regions to be heated independently; particularly useful with large diameter items or where there are variations in thickness.

Temperature gradients may be axial (along the length of a pipe or vessel) and through thickness. The through thickness temperature gradient is caused by heat losses from the internal surface and is a function of both thickness and internal diameter, the larger the diameter, the greater the effect of radiation and convection losses. Both the width of the soak band and the temperature achieved can be substantially less than that on the outside of the pipe or tube. Insulation on the inner surface will reduce the temperature/width differential but may not be possible on small diameter tubes or pipework systems. This through thickness gradient is one of the reasons that specifications and codes require the soak or heated band to be a minimum width, generally related in some way to the thickness of the component.

As mentioned above, there are rules in the application codes concerning the size of the heated area, normally related to the thickness. In a circular component such as a pipe butt weld or a pressure vessel circumferential seam the width of the band is easy to calculate. ASME VIII for instance requires the soak band width to be twice the thickness of the weld or 50.8mm either side of the weld, whichever is the lesser.

ASME B31.3 requires the soak band width to be the weld width plus 25.4mm either side of the weld. BS EN 13445 does not specify a soak band width but instead specifies a heated band width of 5√Rt centred on the weld and where R = component inside radius and t = component thickness. There are no requirements in the ASME codes regarding heated band width. A very approximate rule of thumb for flat plate is that the heated band should be a minimum of twice the length of the weld although practical considerations may prevent achieving this ideal.

There are no requirements, in any code or specification, on the width of the thermally insulated band although BS EN 13445 recommends 10√Re. It is essential that the relevant specification is referred to for specific guidance on what is required and it is worth remembering that the specification requirements on soak or heated band widths are minima and very little is lost by ensuring the specified dimensions are comfortably exceeded.

What is an acceptable axial temperature gradient? Again, there is little advice in the codes and specifications. It is generally assumed that if the temperature at the edge of the heated band is above half that of the soak temperature then the temperature gradient will not be harmful. During heating and cooling BS EN 13445 specifies a maximum temperature difference of 150°C in 4500mm below 450°C (1°C in 3mm) and 1000C in 4500mm above 4500C (1°C in 4.5mm).

To ensure that gradients and temperatures are controlled within acceptable limits sufficient thermocouples need to be attached to provide both temperature control and recording. For small diameter tubes, eg less than 100mm diameter, one control zone and one recording thermocouple are regarded as sufficient; between 100-200mm one control zone and one recording thermocouple at each of the 12 o’clock and 6 o’clock positions; above 250mm diameter one control zone and one recording thermocouple at each 900 quadrant, 12, 3, 6 and 9 o’clock, are suggested.

These thermocouples should be placed on the centre line of the weld. Thermocouples will also be needed at the edge of the soak band and the edge of the heated band. Ideally, thermocouples should also be placed on

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the opposite surface to the heating elements to ensure that the correct through thickness temperature has been achieved although this is rarely possible on pipe systems.  It is advisable to double up on the thermocouples to cope with the possibility of a thermocouple failure.

Thermocouples use a hot and a cold junction to measure the temperature, the hot junction being attached to the component, the cold junction within the temperature recorder. For accurate temperature measurement the hot junction must obviously be at the temperature of the component. Errors can be introduced if the junction is not firmly attached, either by capacitor discharge (CD) welding, by mechanically fixing the wires to the component or by overheating of the thermocouple junction.

CD welding of the thermocouple wires gives the most accurate results, particularly if the two wires are separated by 3-4mm. Mechanically attached wires will probably need to be insulated by covering the junction with heat resistant putty to prevent overheating of the thermocouple by the overlying heater. If the wire covering is stripped back then the bare wires also need to be insulated. It is advisable to specify the positions of the thermocouples on a drawing and to include these within a formal written heat treatment procedure document that covers both the specification and best practice requirements.

For more information, please contact us.

Fig 1. Induction PWHT of Pipework

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Fig 2 Schematic of Temperature bands within a local PWHT (Reproduced with permission of the American Welding Society (AWS), Miami, Florida, USA)

Heat Treatment of welded joints Part 4 - Precipitation or age hardening

Job Knowledge

There are several methods that may be used to increase the strength of a metal; alloying, quenching of steel, work hardening, and one very specific form of heat treatment, that of precipitation or age hardening (the two terms are synonymous). Many ferrous and non-ferrous alloys are capable of being age hardened and, as the name suggests, this method of increasing the strength relies upon the formation of precipitates. To achieve the optimum combination of mechanical properties the heat treatment cycles must be very closely controlled.

Unfortunately, to understand how the precipitates affect the mechanical properties it is necessary to introduce some fundamental metallurgy.

The precipitation hardening mechanism requires the solubility of the alloying element, the solute, in the metal, the solvent, to increase as the temperature increases as shown in the phase diagram in Figure 1 where the solvus line shows decreasing solubility of alloying element B in the solvent A as the temperature falls.

An analogy is that of salt in water; as the temperature increases more salt can be dissolved but the converse happens as the solution is allowed to cool when salt crystals begin to form or precipitate.

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The same process occurs in suitably alloyed metals except that the processes of dissolving and precipitating take place in the solid and are hence much slower as atoms find it more difficult to move in a solid than a liquid solution.

A consequence of this is that once the precipitates have been dissolved by taking the metal alloy to a sufficiently high temperature, ie above the solvus line, they can be prevented from re-forming by rapid cooling or quenching.

This heat treatment is known as solution heat treatment and is carried out to form an unstable super-cooled solid solution which, if reheated to a lower ageing or precipitation hardening temperature, will begin to re-form the precipitates, these growing in size as the heat treatment proceeds. A schematic of such a heat treatment cycle is also given in Figure 1 for alloy N.

Figure 1: Phase diagram showing decreasing solubility of B in A and heat treatment cycle

In the solution heat treated metal the atoms of the alloying element, the solute, are randomly distributed throughout the matrix but once the temperature is raised the precipitates begin to form by a nucleation and growth process. At relatively low temperatures and in a short timescale the solute atoms begin to cluster together to form extremely small and very finely dispersed precipitates known as Guinier-Preston (GP) zones, named after the two metallurgists who first identified them. The GP zones are so small that they are not visible using normal optical microscopes but can be seen using electron microscopy at magnifications of around x100,000.

The GP zones are described as coherent, in other words they have the same crystal structure as the solvent metal. However, they distort the crystal lattice, the framework on which the atoms are positioned. This makes it more difficult for dislocations to move through the lattice and it is dislocation movement that enables metal to deform; tensile strength and hardness therefore increase but ductility and toughness decrease. As the ageing treatment continues or the temperature is raised the tensile strength also continues to increase as the precipitates grow and coarsen whilst still remaining coherent. At some point, however, the precipitates begin to lose their coherency; they become incoherent, forming separate particles within the metal with a different crystal structure from the solvent and at this stage they become visible using an optical microscope.

Just before this point is reached is when the alloy has the very highest tensile strength. As these incoherent particles form and grow in size the tensile strength progressively decreases. The alloy then is said to be

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overaged although the precipitates still contribute towards the tensile strength of the alloy. The high strength low alloy (HSLA) steels are a good example of this where incoherent, overaged precipitates are used to give a substantial increase in the tensile strength.

In order to achieve the best combination of properties the precipitates need to be evenly distributed throughout the grains of the alloy and of an optimum size. The ageing temperature and/or time can obviously be changed to tailor the distribution and size of the precipitates; longer times and/or higher temperatures generally result in a reduction in strength but an increase in ductility, an overaged structure giving the lowest tensile strength but the highest ductility.

Typical heat treatment times and temperatures of a range of different alloys are given in Table 1. The ferritic and nickel based alloys are generally used in the overaged condition in order to ensure a reasonable degree of ductility. It can be seen that with some alloys, eg 17/4PH stainless steel, the precipitation mechanism is sufficiently sluggish that the component can be cooled in still air or, as with the A286 stainless steel, long ageing times are required.

On the other hand the aluminium-copper alloy 2219 is capable of ageing at room temperature if left for a couple of days. Some of the 6000(Al-Si-Mg) and series 7000 (Al-Zn-Mg) alloys will similarly age at ambient temperature. This is known as natural ageing; – aging at an elevated temperature is known as artificial ageing.

Table 1: Typical ageing heat treatments and properties of a range of age hardening alloys

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The close control of heat treatment times, temperatures and cooling rates is therefore essential if the required properties are to be obtained. For the solution treatment of aluminium alloys a salt bath is frequently used, artificial ageing taking place in a forced air circulation furnace. Illustrated in Figure 2 is the effect of varying the time and temperature on the ultimate tensile strength of an Al-4%Cu alloy such as alloy 2025 where it can be seen that a difference as small as 40OC in the ageing temperature can have a major effect on the strength. The higher temperature needed by the nickel and ferrous alloys generally requires the use of gas fired or electrical furnaces with sufficient thermocouples to ensure the correct temperatures are consistently achieved throughout the component.

Figure 2: Effect of varying ageing times and temperatures on tensile strength

A comparison of ISO 15614 Part 1 and ASME IX

Job Knowledge

Job knowledge

The question is sometimes asked ‘Can I use our existing welding procedure qualifications?’ where the qualification specification required by the contract is one that has not previously been used by the organisation. This is particularly relevant when substantial costs and/or delays will be incurred if re-qualification of the welding procedures is necessary. The two most frequently encountered specifications are ISO 15614 Part 1 and ASME IX and whilst these are written with the same purpose (that of giving assurance that a welding procedure will provide the desired joint properties)  there are major differences between the two specifications that mean that they are not equivalent. It will not be possible in this short article to cover every welding variable and its range of approval in the two specifications. Where compliance is required then reference MUST be made to the appropriate specification.

With respect to ASME IX the specification requirements can be applied in two ways; ASME intent and ASME stamp. If the welded item is to be ASME stamped this can only be done by a manufacturer who has a quality system accredited by ASME and who holds an appropriate stamp, N stamp for nuclear components, U for unfired pressure vessels, S for power boilers etc. All the requirements of the ASME specifications

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MUST be complied with, even to the extent of dimensions of the mechanical test pieces and the calibration of testing equipment.

ASME intent is used where the item is not to be code stamped but is perhaps only designed to the relevant ASME code and some flexibility is possible with respect to the manufacturing aspects of specification compliance. Such flexibility may allow the manufacturer to submit to the client or inspecting authority procedure qualification records (PQR) to ISO 15614 Part 1 for approval that can be shown to be technically equivalent to an ASME PQR.

ASME IX covers the qualification of welders and welding operators, welding procedures, brazing operatives and brazing procedures for the complete range of ferrous and non-ferrous engineering metals (steels, copper, nickel, aluminium, titanium and zirconium alloys) and oxy-gas, arc, power beam, resistance and solid phase welding processes. ISO 15614 Pt1 covers the welding procedure qualification of arc and gas welds in steel and nickel alloys only. Other alloys and joining processes are covered by additional specifications within the ISO 15614 series.

Both specifications identify essential variable (although ISO 15614 Pt1 does not describe them as such) to each of which is assigned a range of approval. A change to an essential variable outside of its range of approval requires the welding procedure to be re-qualified. ASME IX in addition identifies supplementary and non-essential variables. Supplementary variables are only invoked when toughness requirements are specified by the application code, eg ASME VIII or ASME B31.3. Non-essential variables, as the name suggests, are those variables that are not regarded as affecting the quality or mechanical properties of the welded joint and comprise such variables as the weld preparation, shield gas flow rate, method of back gouging, shield gas nozzle size etc. Although these variables are non-essential it is a requirement that they should be referenced on the welding procedure. It is therefore NOT acceptable to use a butt welding procedure to specify how a fillet weld should be made.

ISO 15614 Pt1 does not identify any variables as non-essential;  where a variable is not regarded as significant it is simply not referenced in the specification. There are several variables in both specifications where there is no range of approval;  the manufacturer, the welding process and the application or deletion of post weld heat treatment (PWHT) for example.

In order to reduce the amount of qualification testing, both specifications group alloys of similar characteristics together. Qualifying the welding of one alloy within the group allows the other alloys within the group to be welded. ASME IX assigns the groups  numbers with steels being numbered P1 to P15F. Any alloy that does not have a P number is regarded as unassigned; a procedure qualification carried out using an unassigned alloy qualifies only that specific designation of alloy. Until recently only alloys that complied with the ASME and/or ASTM material specifications and/or had a UNS number were assigned P numbers. However, a limited number of EN, Canadian, Chinese and Japanese alloys have now been introduced into the list of assigned alloys.

ISO 15614 Pt. 1 also groups steel and nickel alloys into families with similar properties but is somewhat less prescriptive than the ASME code in that, provided alloys have similar chemical compositions and mechanical properties, the material specification is not relevant – for example a plain carbon steel with less than 0.25%C and a minimum specified yield strength less than 460MPa  falls into Group 1 irrespective of whether or not it is a pressure vessel or structural steel or supplied in accordance with EN or ASTM material specifications. To determine into which group the alloy falls reference should be made to ISO/TR 15608, the specification that lists both ferrous and non-ferrous alloys and assigns them a group number.

Other significant differences between the two specifications with respect to the arc welding processes are :-

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ASME IX requires only tensile and bend tests to qualify a butt weld. ISO 15614 Pt1 requires a far more extensive test programme of visual inspection, radiography or ultrasonic examination, surface crack detection, tensile and bend tests and macro-examination. In certain circumstances Charpy-V impact tests and hardness surveys are also required. 

ASME IX specifies that the tensile strength of the cross joint tensile specimen shall be at least that of the minimum specified for the parent metal and that bend test coupons should have no discontinuity greater than 3mm. ISO 15614 Pt1 has identical requirements for these mechanical tests but in addition specifies an acceptance standard for the non-destructive testing; impact test results, when required, that match the parent material toughness and hardness limits when hardness testing is required.

ISO 15614 Pt 1 requires Charpy-V impact testing for steels over 12mm thick when the material specification requires it. ASME requires impact testing only when specified in the application standard. This requirement makes heat input a supplementary essential variable in ASME IX but an essential variable in ISO 15614 Pt1.

Hardness testing is required by ISO 15614 Pt1 for all ferritic steels with a specified minimum yield strength greater than 275MPa. A maximum hardness for joints in either the as-welded of PWHT’d condition is specified. ASME IX does not require hardness testing.  

ASME IX allows a reduction in preheat of 55OC before requalification is required. ISO 15614 Pt1 does not permit any reduction in preheat from that used in the qualification test.

ASME allows the maximum interpass temperature to be 55OC above that measured in the qualification test. ISO 15614 Pt 1 permits no such increase.

ASME IX requires pressure containing fillet welds to be qualified by a butt weld procedure qualification test. Non-pressure retaining fillet welds may be qualified by a fillet weld test only. ISO 15614 Pt1 requires a fillet weld to be qualified by a butt weld when mechanical properties “.... are relevant to the application...” i.e when it is a load carrying fillet weld. In addition, whilst a butt weld will qualify a fillet weld “....fillet weld tests shall be required where this is the predominant form of production welding...” i.e. an ISO compliant welding procedure where the majority of the welding is of load carrying fillet welds must reference both a butt weld and a fillet weld procedure qualification.

Weld metal transfer mode, where relevant, is an essential variable in both ISO 15614 Pt1 and ASME IX but the current type is an essential variable in ISO 15614 Pt1 and a supplementary essential variable in ASME IX.

A change from manual to automatic welding is an essential variable in ISO 15614 Pt1 but a non-essential variable in ASME IX.

Whilst there are several other variables in the two specifications that have substantially different ranges of approval there are many that have ranges that are very similar – material thickness being but one example. 

This article has highlighted some of the significant differences but to ensure that the welding procedure and its supporting procedure qualification record are compliant the specifications must be referred to. The answer to the question posed at the start of this article is therefore – it depends upon what you can persuade the client and inspecting authority to accept!

Power source characteristics

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Fig. 1. Static arc characteristic

The prime objective of an arc welding power source is to deliver controllable welding current at a voltage demanded by the welding process. The arc welding processes have different requirements with respect to the controls necessary to give the required welding conditions and these in their turn influence the design of the power source. In order to understand how the requirements of the processes affect the design of the power source it is necessary to understand the interaction of the power source and the arc characteristics.

If the voltage of a welding arc at varying arc lengths is plotted against the welding current the curves illustrated in Fig. 1 are obtained. The highest voltage is the open circuit voltage of the power source. Once the arc is struck the voltage rapidly falls as the gases in the arc gap become ionised and electrically conductive, the electrode heats up and the size of the arc column increases. The welding current increases as the voltage falls until a point is reached at which time the voltage/current relationship becomes linear and begins to follow Ohms Law. What is important to note from Fig. 1 is that as the arc length changes both the voltage and welding current also change – a longer arc giving higher voltage but with a corresponding drop in welding current and vice versa. This characteristic of the welding arc affects the design of the power source since large changes in welding current in manual metallic arc (MMA) and TIG welding is undesirable but is essential for the MIG/MAG and flux cored arc welding processes.

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Fig 2 Constant current power source characteristic

MMA, TIG and submerged arc power sources are therefore designed with what is known as a drooping output or constant current static characteristic, MIG/MAG and FCAW power sources with a flat or constant voltage static characteristic. On most power sources the slope of the characteristic can be changed either to flatten or make steeper the curves shown in Fig 2 and Fig. 3

Fig 2 shows drooping or constant current power source static characteristics, such as would be used for the MMA or TIG process, superimposed on the arc characteristic curves. When manual welding is taking place the arc length is continually changing as the welder cannot maintain a constant arc length. With a constant current power source as the arc length changes due to the welder’s manipulation of the welding torch there is only a small change in the welding current – the steeper the curve the smaller the change in current so there will be no current surges and a stable welding condition is achieved. Since it is primarily the welding current that determines such features as the penetration and electrode consumption this means that the arc length is less critical, making the welder’s task easier in achieving sound defect free welds. Typically, a ±5volt change would result in around a ±8 amp change at 150amp welding current.

In some situations – for example when welding in the overhead position or when the welder is faced with variable root gaps - it is an advantage if the welder has rather more control over deposition rates by enabling him to vary the rate by changing the arc length. In such a situation a flatter power source characteristic will be of benefit.

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Submerged arc welding also uses a drooping characteristic power source where the welding current and the electrode feed rate are matched to the rate at which the wire is melted and transferred across the arc and into the weld pool – the “burn-off rate”. This matching of parameters is carried out by a monitoring system which uses the arc voltage to control the electrode feed speed – if the arc length/voltage increases the wire feed speed is increased to restore equilibrium.The constant voltage power source characteristic is illustrated in Fig. 3. This shows that as the arc length and hence the voltage changes there is a large change in the welding current – as the arc lengthens the welding current falls, as the arc shortens the current increases.

Fig. 3 Constant voltage power source characteristic

With MIG/MAG and FCAW power sources the welding current is controlled by the wire feed speed, the welding current determining the rate at which the welding wire is melted and transferred across the arc and into the weld pool – the “burn-off” rate. Therefore, as the current decreases the burn-off rate also falls, less wire is melted and the wire tip approaches the weld pool. In doing so, the voltage decreases, the welding current and hence the burn-off rate increase. Since the wire feed speed is constant there is a surplus of burn-off over wire feed such that the desired arc length, voltage and current are re-established. The converse also occurs – a shortening of the arc causes a reduction in voltage, the current rises, the burn-off rate increases, causing the arc to lengthen, the voltage to increase and the welding current to fall until the pre-set welding conditions are re-established. Again, a typical figure for the change in welding current for a constant voltage power source would be in the region of ±40amps for a change in arc length of ±5volts. This feature gives us what is known as a “self-adjusting arc” where changes in arc length, voltage and current are automatically returned to the required values, producing stable welding conditions. This makes the welder’s task somewhat easier when compared with MMA or TIG welding. Although in principle it may be possible to

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use a constant voltage characteristic power source for MMA welding it is far more difficult for the welder to judge burn-off rate than arc length so arc instability results and the method is not practicable.

In addition to this voltage control of the welding arc the speed at which the power source responds to short circuiting is important - this is known as the power source dynamic characteristic. Short circuits occur during arc striking and in MIG/MAG welding during dip transfer. As the voltage drops to zero when a short circuit occurs the current rises. If this increase in the current is fast and uncontrolled then the electrode tip blows like an electrical fuse resulting in excessive spatter – too slow a rise and the electrode may stub into the weld pool and extinguish the arc. This is not too significant when using the MMA process since the maximum current at zero voltage is controlled by the slope of the static characteristic curve and the welder can easily establish an arc gap. It is, however, important in the MIG/MAG process where a flat static characteristic power source is used and the current could rise to an extremely high value, in particular when welding in the dip transfer or short circuiting condition.

An electrical component called an inductor is therefore introduced into the power source electrical circuit. This device opposes changes in the welding current and hence slows the rate at which the current increases during a short circuit. The inductance is variable and can be adjusted to give a stable condition as shown in Fig. 4. Inductance in the welding circuit also results in fewer short circuits per second and a longer arc-on time - this gives a smoother better shaped weld bead. Too much inductance, however, may result in such a slow rise in the welding current that there is insufficient time for the arc to re-establish and melt the wire tip so that the welding wire then stubs into the weld pool. Inductance during spray transfer is also helpful in providing a better and less violent arc start. 

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Non-destructive Examination (NDE) Part 1 Liquid Penetrant and Magnetic Particle Inspection

Job Knowledge

Fig. 1 Principles of LPI

In achieving high quality defect free welds there is no substitute for experienced and qualified welders and competent supervision. However, no matter how skilled the welder, during the process of depositing weld metal, imperfections of various types may be formed. It is therefore necessary to have methods of ensuring that the weld is of an acceptable quality, hence the development of a range of non-destructive inspection techniques capable of both detecting and sizing buried and surface breaking imperfections, enabling a decision to be made regarding acceptance or otherwise.

Note that both the ISO specifications and the ASME codes differentiate between an ‘imperfection’ (ISO) or a ‘discontinuity’ (ASME) and a ‘defect’. It is accepted that all welds contain features or imperfections but it is only when an imperfection exceeds the relevant acceptance standard does an imperfection become a defect – as ASME states “..... this term designates rejectability...”    

Liquid penetrant (LPI) and magnetic particle (MPI) inspection techniques are methods that supplement visual inspection, revealing defects such as fine cracks or micro-porosity that would be invisible or difficult to detect by the naked eye. 

LPI is a simple, cheap and easily portable inspection method that requires no equipment apart from spray cans. It can detect surface breaking imperfections only and relies on a coloured or fluorescent dye, sprayed on the surface and penetrating the imperfection. About15 minutes is generally specified to enable the dye to penetrate any very fine imperfections. After cleaning the excess the dye is drawn to the surface by spraying on a developer in the case of the colour contrast dye or exposing the surface to ultra-violet light in the case of a fluorescent dye, the imperfection being revealed by the dye staining the developer or fluorescing, as shown schematically in Fig. 1 and in Fig. 2, a liquation crack in an aluminium alloy. 

The fluorescent dye gives greater sensitivity than the colour contrast dye and does not require the use of a developer. It does however require the use of an ultra-violet light source and preferably a darkened room which makes it a less portable inspection method than the contrast dye technique.

The dye used as a penetrant must be able to penetrate tight cracks but must not be capable of being removed from more open imperfections during the cleaning operation carried out prior to applying the developer.

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Fig. 2 LPI test result – cracking in the HAZ of an aluminium alloy weld

Careful surface preparation and thorough cleaning of the item before applying the penetrant is important. Swabbing with or immersion of the item in a proprietary degreasant is generally sufficient – cleaning in an ultrasonic bath is the best method but can be used only for small portable components. Grinding or wire brushing, particularly of materials such as copper and aluminium alloys, should be avoided if possible as such cleaning methods can smear over imperfections, making them undetectable. If such is the case an acid etch may be required to remove the smeared metal and enable the dye to penetrate the imperfectionInspection in positions other than flat can be a problem but penetrants are available with a jelly like consistency that can be used in the vertical and overhead positions. It is possible to automate the process with small components loaded into baskets and processed on a conveyor line. Fluorescent dyes are better than contrast dyes in this application due to their greater sensitivity.

Although a simple inspection process to use, interpretation can be a problem if the surface is naturally rough – coarsely ground or rough machined for example - or contains acceptable geometric features that trap the dye. Training of operatives to recognise genuine and spurious indications is therefore essential. Other limitations are that it can be used at room temperature only and it is not possible to indefinitely retest components as the imperfection becomes filled with dry dye. Health and safety may also be an issue with irritation of unprotected skin and fumes from some of the cleaning and solvent chemicals, particularly when the process is used in confined spaces. 

The process can however be used to inspect both ferrous and non-ferrous metals, large areas can be examined very quickly and it can be used on components with complex geometry.

Magnetic particle inspection (MPI) is also a simple-to-use inspection method but, as the name suggests, is limited in use to magnetic materials- in other words ferritic (NOT austenitic) steels. The basic principle is that the component is magnetised, producing a flux within the metal as shown in Fig 3 An imperfection is non-magnetic and therefore cuts the lines of flux producing a leakage field around the imperfection –a localised “magnet”. A magnetic powder sprayed or dusted onto the surface will be attracted to this “magnet” forming a line of powder. The strength of the magnetising current should be specified in a written

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examination procedure and the adequacy of the magnetic field verified by the equipment being capable of lifting a specified weight.  

[ Zoom ]

Fig. 3 Principles of MPI

Conventionally the item to be inspected is spray painted with a thin coat of rapid drying white paint, the magnetic field is applied and a black magnetic ink is sprayed onto the surface, forming a black indication against the white background. The maximum imperfection sensitivity is when the imperfection cuts the magnetic flux at 90O. In order that both longitudinal and transverse oriented imperfections are detected, examination of a weld must therefore be carried out with the magnet applied twice at 90O along and across the weld .  

Magnetisation may be by prods, electromagnetic yokes as shown in Fig 4, or permanent magnets. Inspection by the use of prods supplied with high amperage low voltage alternating or direct current is often used in the workshop, the local magnetisation being achieved by two prods connected to a transformer or transformer/rectifier. The prods are pressed onto the metal surface, a trigger pulled to initiate a current in the component and the magnetic ink applied. This is generally a two man operation. Care must be taken not to initiate the current before the prods are in firm contact with the surface as arcing between the prod tips and the component can occur, resulting in a feature similar to a welding arc strike. The use of rectified half wave single phase direct current has an advantage over alternating current and the yoke in that the process is capable of detecting imperfections up to perhaps 1mm below the surface, depending on their size and orientation. 

[ Zoom ]

Fig. 4 MPI using an AC yoke

The yoke method, illustrated in Fig 4, has several advantages over the prod method. The equipment is relatively small and lightweight; can be battery powered and is readily portable, making it ideal for site inspection. The yoke can be operated in one hand and the magnetic ink sprayed on from the other making this a one man operation. In addition no electrical current is transferred into the component.

Spurious indications can be produced where there is a difference in magnetic properties within the joint, perhaps in the HAZ where it may be mistaken for unacceptable undercut. Two metals of different magnetic characteristics when joined together can give a well defined indication suggesting the presence of a crack – a careful dressing of the surface followed by LPI is helpful in deciding whether or not the indication is genuine. Ferritic/austenitic dissimilar metal joints cannot be MPI’d. Residual magnetism can also cause

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problems in interpretation. As with LPI rough uneven surfaces may also give rise to spurious indications. 

Neither technique gives a permanent record of the inspection but where this is necessary photographs of the affected area are very useful. Conveniently positioned reference markers and a scale are helpful for the accurate recording of the size and position of indications, particularly if repairs are required. It is also possible to transfer the indication onto transparent sticky tape by carefully pressing the tape onto the surface and then applying this to a sheet of white paper.Eddy Current Testing

Job Knowledge

Eddy current testing is an inspection method that can be used for a variety of purposes including the detection of cracks and corrosion, material and coating thickness measurement, material identification and, in certain materials, heat treatment condition. The process relies upon a material characteristic known as electro-magnetic induction. When an alternating current is passed through a conductor – a copper coil for example – an alternating magnetic field is developed around the coil, the field expanding and contracting as the alternating current rises and falls. If the coil is then brought close to another electrical conductor the fluctuating magnetic field surrounding the coil permeates the material and induces a circulating or eddy current to flow in the conductor. This eddy current, in its turn, develops its own magnetic field. This ‘secondary’ magnetic field opposes the ‘primary’ magnetic field and thus affects the current and voltage flowing in the coil. Any changes in the conductivity of the material being examined such as near surface defects or differences in thickness will affect the magnitude of the eddy current and this change can be detected using either the primary coil or a second detector coil.  This forms the basis of the eddy current inspection technique. 

As with any inspection method there are both advantages and disadvantages to eddy current testing. The method can be used only on conductive materials and, although all metals can be inspected, the depth of penetration of the eddy currents varies. Eddy current density is higher and defect sensitivity greatest at the surface and decreases with depth, the rate of the decrease depending on the “conductivity” and “permeability” of the metal. The conductivity of the material affects the depth of penetration with a greater flow of eddy current at the surface in high conductivity metals and a subsequent decrease in penetration in metals such as copper and aluminium.

Permeability is the ease with which a material can be magnetised. The greater the permeability the smaller will be the depth of penetration. ’Non-magnetic’ metals such as austenitic stainless steels, aluminium and copper have very low permeability whereas the ferritic steels have a magnetic permeability several hundred times greater. 

The depth of penetration may be varied by changing the frequency of the alternation current – the lower the frequency the greater is the depth of penetration. Unfortunately, as the frequency is decreased to give this greater penetration the defect detection sensitivity is also reduced. There is therefore, for each test, an optimum frequency to give the required depth of penetration and sensitivity.

A parameter known as the “standard depth of penetration”, taken as the depth at which the eddy current value has reduced to 37% of that at the surface, can be calculated from the magnetic permeability, the metal’s conductivity and the frequency of the alternating current in the probe. The standard depth of penetration is generally regarded as the criterion by which the efficiency of detection can be judged, although changes in the eddy current can be detected at depths of up to three times this figure. A simple calculation may be used to select the optimum probe frequency.     

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Fig 1 Pen type probe being used to examine bolt holes for cracks.

For any particular inspection the accuracy of the measurement of defect size, material thickness, heat treatment condition etc. is largely determined by the design of the coil (or coils) used in the examination whilst detection capability is also determined by material properties and the equipment characteristics. The selection of the probe is therefore critical for accurate results. 

Some inspections involve sweeping through multiple frequencies to optimize results, or inspection with multiple coils to obtain the best resolution and penetration required to detect all possible flaws. It is always important to select the right probe for each application in order to optimize test performance. 

 The eddy current operator is therefore faced with a material whose conductivity and permeability are physical properties and outside of the operator’s control. The parameters that can be selected are probe size, probe type and frequency of the alternating current, the selection depending upon the test requirements i.e crack detection, corrosion depth, coating thickness, heat treatment condition etc. Some equipment is designed to operate using multiple frequencies or with multiple probes in order to optimize the test performance and achieve the best detection performance and depth of penetration. The results are displayed either as a digital read-out for the more simple examinations such as thickness measurements or displayed on an oscilloscope screen as an X-Y display of resistance versus the inductive reactance. This gives a characteristic curve, the shape and size of which can be used to detect and size a defect as illustrated in Fig.2 to determine heat treatment condition or, as a quick sorting test, to establish the type of alloy.

Fig. 2 Schematic of Screen Display

In addition to selecting the optimum frequency the size of the probe can be varied – a large diameter coil will inspect a larger volume of metal and therefore reduce the inspection time – a small diameter probe, however, is more sensitive and better suited to detecting small flaws. The large diameter probes are often used for the detection of large sub-surface flaws in castings and forgings and for the detection of corrosion; the small diameter pencil type probes for detecting cracks. Weld examination requires special probes to reduce noise from the permeability change across a weld. 

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As mentioned earlier eddy current testing can be used for a variety of inspection tasks. Chief amongst these is the inspection of welded joints using pencil probes as a replacement for the more conventional magnetic particle or liquid penetrant inspection techniques. A major advantage is that the process may be used underwater and can be used to scan welds through paint and other coatings. With respect to detection of linear defects such as cracks and lacks of fusion the defect should break the lines of the eddy currents ideally at right angles – as with magnetic particle inspection defects parallel to the eddy currents are likely to remain undetected. It is important therefore that the weld is scanned in the correct direction. Cracks as small as 0.5mm deep and 5mm in length are capable of being detected.

Encircling coils are used for automated in-line tube inspection with welded tube lines using localised probes for weld examination. 

By measuring the conductivity of a metal it is possible to identify and sort both ferrous and non-ferrous metal and with certain alloys - in particular the aluminium alloys - it is also possible to establish the heat treatment condition. Low frequency probes are used to detect generalised corrosion, particularly in the aerospace industry for the examination of aircraft skins. Specially designed “bobbin” probes can be used to inspect the bore of tubes in service for signs of pitting or corrosion and there are also probes specially designed to examine the bores of bolt holes for cracks. 

Measuring the proximity of a component to the probe can also be used to determine coating thickness provided the coating is non-conductive. The “lift-off”, the distance of the probe tip from the conductive surface, causes a change in eddy current flow which is measurable. 

All of the systems must be calibrated using appropriate reference standards – as for any NDT method, this is an essential part of any eddy current examination procedure. The calibration blocks must be of the same material, heat treatment condition, shape and size of the item to be tested. For defect detection the calibration block contains artificial defects simulating defects; for corrosion detection a calibration block of different thicknesses is used. 

The eddy current method requires more skill on the part of the operator than, say, MPI and penetrant inspection – it goes without saying that operator training is essential.Radiography

Job Knowledge

Fig. 1. Principles of radiography

The previous two Connect articles dealt with the defect detection techniques of MPI, LPI and eddy current testing. These methods are capable of detecting surface or very near surface imperfections and some process is therefore required to enable buried imperfections to be reliably detected – a so-called volumetric detection method. The first such method to be used in manufacturing to determine the quality of fabricated components is radiography. X-rays generated from a cathode ray tube were discovered in the late 1890’s, soon followed by the discovery of gamma radiation from radioactive isotopes. 

The ability of this radiation to penetrate both living and inert objects with the amount of transmitted radiation depending on the density, thickness, and atomic number of the object was eventually understood. 

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The transmitted radiation was found to produce images on photographic film and these findings resulted in the development of industrial radiography as illustrated in Fig. 1. 

Fig. 2 Schematic of an X-ray tube

X-radiation is produced from a vacuum tube - this contains a cathode which produces a beam of electrons when an electric current is passed, the higher the voltage the more intense is the stream of electrons and the deeper the penetration – see Table 1. Industrial X-ray tubes use voltages between 20kV and some 450kV but up to 30MV in linear accelerator and betatron equipment. The stream of electrons impacts an anode, some 1% of the energy being emitted as a beam of X-rays, the remainder being released as heat, requiring the anode to be cooled, often with water or oil. A schematic of an X-ray tube is shown as Fig. 2. The anode can be designed to provide a beam as illustrated in Fig. 2 or a 360O panoramic beam, typically used for pipe butt weld inspection with the X-ray source inside the pipe.

kVMaximum Thickness (mm)

Steel Aluminium

50 1 12

100 20 60

150 30 80

200 50 100

300 80 150

400 100 200

2000(2MV) 200 500

25000(25MV) 500 1400

Table 1 Applied kV and maximum thickness

Gamma radiation is naturally occurring and is produced by the decay of a radioactive isotope which, when it decays, emits three types of radiation, alpha, beta and gamma. Alpha and beta radiation are short range and very easily absorbed – gamma radiation, however, is very energetic, typically over 100keV, and can easily pass through a metal, the thickness that can be penetrated depending on the type and size of isotope. The isotopes are stored in shielded containers from which they can be wound to expose the isotope, as shown in Fig 3.

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Fig 3. Typical arrangement of gamma ray NDE equipment

Each isotope has a “half life”- a length of time by which half (50%) of the radioactive isotope has decayed into a stable element – two half lives means the source has only 25% of its original strength, three half lives 12.5%. As the source decays and becomes less energetic the length of the exposure time must be increased to achieve the same density of image on the radiographic film. There is thus a point at which use of the gamma ray source is discontinued and replaced with a fresh isotope. In addition, because the gamma radiation emitted by an isotope cannot be varied in quality there is a range of material thickness for which each source will give acceptable results. The commonest isotope in regular use is iridium192 with cobalt60 being used for very thick components.

Isotope Half Life Typical Thickness Range – steel  (mm)

Iridium 192 74 days 10 - 50

Cobalt 60 5.26 years 25 - 200

Ytterbium 169 32 days < 10

Thulium 170 128 days < 10

Selenium 75 119 days 5 - 20

Caesium 137 30 years 20 - 80

Table 2 Half life and typical thickness range of industrial isotopes.

The film is a fine grained photographic film with a coating of a light sensitive emulsion on both sides which is loaded in a darkroom into a flexible cassette. The cassette contains an intensifying screen, a card mounted thin lead foil, 0.05 mm to 0.5mm thick, held in close contact with the film. This screen absorbs stray radiation and emits electrons that enable the exposure time to be substantially reduced without affecting the radiographic quality. 

The quality of a radiograph is assessed using three factors – density, contrast and definition or sharpness of the image. Most specifications require film densities in the range 1.8 to 2.5  (film density = 1, 1/10th of light is transmitted: film density = 2, 1/100th of light is transmitted). The density of the film can be measured using a densitometer; films outside of the specified range of density would be rejected. The density is affected by the exposure time and metal thickness – this can make the radiography of components with dissimilar thicknesses problematic – thin sections being over-exposed, thick sections under-exposed. 

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Fig 4. Formation of a penumbra

The contrast is determined by the differences in absorption between the metal and the defect and by the type of film used. The speed of the film in particular affects the  contrast of the image. Most  weld defects are less absorbing than the surrounding metal – slag, porosity, lack of fusion or penetration etc therefore appear dark against a lighter background. The only weld defects that are more absorbingr than the parent metal are tungsten inclusions that appear as bright white specks on the film.

Sharpness of the image is a function of a number of factors, a radiograph with poor sharpness being somewhat similar to an out-of-focus photograph with fine details being blurred or not visible. Film with a very fine grain size is preferred for high quality radiography, being capable of resolving fine details. There is a geometric unsharpness caused by the size of the radiation source, known as the “focal spot”. A large diameter source will cause a penumbra which means that the edges of an image become blurred as shown in Fig. 4. The further away from the item being radiographed than the less obvious is this effect – unfortunately the further away the source is from the object then the longer is the exposure time – twice the distance quadruples the time. To minimise the penumbra and increase sharpness the source should be the smallest diameter possible; the film should be as close to the back of the object as possible; the source should be as far from the object as possible, bearing in mind the lengthening of exposure time; fine grained film should be used. Additional unsharpness is caused by the release of electrons within the film emulsion that darken the adjacent area.

To ensure that radiographs are of an acceptable quality with respect to image sharpness, contrast etc it is a requirement that an image quality indicator (IQI) is used as can be seen in Fig. 1. This topic will be covered in the next article, as will the techniques for radiography of a variety of joint types and component configurations.Radiography Part 2

Job Knowledge

The previous article dealt mostly with the basic principles of radiography – this part will cover the methods of ensuring that a radiograph is of an acceptable quality and capable of showing relevant imperfections. As mentioned in the previous article the quality of a radiograph is assessed using three factors: density, contrast and definition or sharpness of the image. Density and contrast have already been covered but there also has to be some method by which the sensitivity (the ability to reveal imperfections) can be measured. To do this devices known as image quality indicators (IQI’s), formerly called, are used. These can be of several forms as illustrated in Fig.1

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Fig. 1. Image Quality Indicators

The wire type is the most frequently used IQI in radiography-by-film. The design of the IQI is given in EN ISO 19232 Part 1 or ASTM E747. Both specifications list a series of IQIs containing six or seven wires of increasing diameter, from 1-8mm, from 10-50mm in length and in a range of metals – iron, nickel, aluminium, magnesium, copper and titanium. The wires are mounted side by side in a flexible plastic sheath which also carries appropriate identification, generally lead letters that will be clearly seen on the radiograph. The IQI is selected with respect to the metal type and the component thickness; the thicker the component the thicker the IQI’s wires. The IQI carries an identification and serial number so that it can be confirmed at a later stage that the correct IQI has been used.

Ideally the IQI is placed on the source side of the component and, in the case of a weld, transversely across the joint although this is not always possible when radiographing pipe and tube butt welds. The sensitivity is taken as the smallest diameter wire that can be seen divided by the component thickness, expressed as a percentage. Most application codes specify a sensitivity of between 2-4% ; this is a maximum, the smaller the figure the greater the sensitivity of the radiograph. Alternatively an actual wire diameter that must be visible is specified. 

The step hole IQI is used less frequently. It is a stepped wedge with a hole drilled in each step, the hole diameter matching the thickness of the step. As with the wire IQI, the material and dimensions of the step wedge are selected to match the application. The diameter of the smallest hole visible on the radiograph determines the sensitivity, this being calculated as hole diameter divided by component thickness expressed as a percentage. The sensitivity measured by the use of a wire IQI is not the same as the sensitivity using a step wedge IQI. 

As with any film, the method of processing will affect the quality of the image. Care must be taken to ensure that there is no light contamination, the processing chemicals are at the correct concentrations and temperatures and that drying the film does not leave marks and stains that would leave spurious indications and would make accurate interpretation difficult.

Interpretation of the radiographs must be undertaken by trained and experienced radiographers. In addition to being fully conversant with radiographic techniques such individuals should also have a comprehensive knowledge of welding processes, joint design and the various imperfections that may occur.  Many application specifications require such individuals to be independently certified to a suitable certification

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scheme such as PCN, administered by the British Institute for NDT or CSWIP, administered by TWI Certification Ltd. Viewing should be carried out in a darkened room, allowing a period of time for the viewer’s eyes to be accustomed to the conditions. The luminance of the viewing screen will need to vary with the density of the radiograph – there is generally a rheostat control (a dimmer switch) on the viewer to enable the luminance to be varied. The light itself should be white and diffuse and there should be as little light as possible leaking around the edges of the radiograph.  

Radiography of flat plates and cylinders large enough to permit entry for placement of the film is a relatively simple operation, as shown in Fig. 1 of the previous article. Lead numbers are placed at fixed intervals along the plate adjacent to the weld or around the circumference of a pipe to enable the position of any imperfections to be accurately located. The weld also carries a unique identification number reproduced on the radiograph by the use of lead figures placed adjacent to the weld.

Radiography of pipes, however, where access to the bore to place the film is not possible presents some problems and terms such as SWSI and DWSI are used as shorthand to identify the various techniques that may be used.

Single wall, single image (SWSI) is a technique whereby the radiographic source is placed inside the pipe by some suitable method, the film wrapped around the outside of the pipe and the exposure made as shown in Fig. 2. This may also be known as a panoramic exposure. The IQI is placed on the outside of the pipe immediately beneath the film. Both X- and gamma-radiography can be used, the source being placed in position by the use of a pre-placed spider or by means of a crawler unit. This method is most commonly used for the inspection of pipelines where the weld can be radiographed in one exposure, making the technique rapid and cost effective.

Fig. 2 Single wall, single image (SWSI) or panoramic radiographic technique

Where access to the bore is not possible or the pipe diameter is too small to permit the use of an internal source then the double wall, single image (DWSI) technique is used. Here the film is placed on the outside of the pipe on the farthest side from the radiographic source, as shown in Fig. 3. The source may be offset slightly to avoid an image of the upper part of the weld to be projected onto the film or directly in line. The source may be close to or a substantial distance from the pipe, the location being a compromise between a less sharp image but short exposure time for a small stand-off and sharper image but longer exposure time for a large stand-off. The need to penetrate two wall thicknesses means that the sensitivity will be poorer than with the single wall single image technique. The technique also requires multiple exposures to enable the complete circumference of the pipe to be examined – specification or contract requirements frequently specify the minimum numbers of exposures to ensure complete coverage and images of an acceptable quality. The technique is generally used on pipes over 80mm in diameter.

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Fig. 3 Double wall, single image

The last technique is double wall, double image (DWDI), generally used only on pipes less than 75-80mm in diameter. By offsetting the source from the weld centre line and using a long source to film distance it is possible to project an image onto the film of both the upper and the lower parts of the weld as shown in Fig. 4. As with the DWSI technique multiple exposures are required to achieve complete coverage.

Fig. 4. Double wall double image radiograph of a pipe butt weld. Note the IQI, identification numbers and position markers.

Part 3 will look at some of the more sophisticated radiographic techniques and the advantages and disadvantages of radiography.

Radiography Part 3

Job Knowledge

The previous two articles dealt with what may be termed conventional radiographic techniques employing either gamma- or X-ray sources and photographic film. The development of electronics, in particular the increase in computing power over the last 20 or 30 years, has enabled what were laboratory based radiographic methods - real time radiography and computerised tomography (the “CT” scan that is perhaps more familiar in a medical context) -to be implemented both on the shop floor and on site.  

The fundamental difference between film radiography and real time radiography is the way in which the radiographic image is handled - the image in real time radiography being produced electronically rather than on film. This means that the image can be viewed, as the name suggests, in real time – an instant result rather than the significant time delay that occurs when radiographic film is carried to a darkroom, developed and dried before viewing can take place. The real time image is viewed on a monitor screen and the results can be stored and transmitted electronically. An additional difference from film radiography is that the image is a positive rather than the negative image of film radiography, denser materials transmitting less radiation and therefore appearing darker and hence voids, slag entrapment, porosity etc show as light areas in a welded joint. 

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The early real time radiographic equipments used a fluorescent screen which interacted with the x-ray radiation to produce an image that was then passed through an image intensifier. This enabled a video film to be produced and the image to be displayed on a monitor screen in real time. Developments in computing power now enable the image to be digitised and enhanced and then analysed. Comparison with a set of pre-programmed parameters e.g. an acceptance standard, enable the process of inspection and acceptance/rejection to be automated. 

Fluorescent screens to a great extent have been replaced by flat or curved screen photodiode arrays based on Si sensors with other alloys e.g. PbI and HgI, being researched in order to improve further the sensitivity of the examination method. Photo-conductors such as selenium or cadmium telluride are also used in what is termed “direct conversion” to give a sharper image than that from the photo-diode panels. The X-ray source has been progressively reduced in size such that the focal spot is now as small as 2 or 3 microns and with micro-focus units to as small as 0.1mm, again resulting in improved image quality. These features also enable the image to be enlarged with magnifications of over 1000 times being available with little or no loss in image quality. 

Real time radiography is employed in applications where a rapid in-production inspection is required. It has found extensive use in the electronics industry for the in-line examination of circuit boards, in the aerospace and automotive industries for the examination of castings and, often using gamma ray sources, in the process industry for the detection of corrosion. By moving the component between the X-ray source and the detector screen seam welded tube and tube butt welds can be inspected.  

Robust, portable manipulating equipment is commercially available for the on-site examination of process pipework and cross country pipelines. Orbiting heads carrying both an X-ray source and, diametrically opposed, the detection screen, are capable of radiographing pipe butt welds with a substantial reduction in time compared with a conventional film radiograph, particularly where multiple exposures are required to give full coverage of the weld. The head rotates around the pipe to give a DWSI (double wall single image) with very high resolution and good contrast. Internal crawlers equipped with a rotating X-ray head and an external detector can be used to provide a SWSI (single wall single image) radiograph. The examination results are instantly available for display on a laptop screen, the images being stored on disc or memory stick. No film processing or darkroom are required resulting in a further reduction in cost. A typical commercially available unit is illustrated in Fig 1.

Fig 1. Portable Real Time Radiographic unit for pipeline inspection. Courtesy Shaw Pipeline Services.

Computed tomography (CT) is a process that uses the techniques of real time radiography to produce three dimensional, rather than two dimensional, images of components. Both the external surfaces and the internal

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structures of an item can be imaged. All materials – metals, plastics, composites, rocks, fossils, the human body – can be scanned, making CT scanning a very powerful investigatory and diagnostic tool. Industrial CT scanning is used in many applications for the internal inspection of components, metrology, and for reverse engineering. 

The earliest CT units utilised an X-ray beam collimated to create a fan shaped beam of radiation which is scanned across the item, the beam being collected at the detector screen. The signal is then manipulated by a powerful software programme, enabling a 3D image to be created. The technique was further improved when the fan shaped beam was replaced by a cone shaped beam. Rotating the part beneath the beam enables a sequence of 2D images to be obtained, generally between 360 (one image per degree) to 3,600 images, depending on the desired resolution, with scanning of the object being completed in as little as a few seconds. The images are subsequently combined using complex software to provide a 3D image of the item, the image containing all of the information relating to both the external and internal surfaces of the component. A typical commercially available 225Kv microfocus unit is shown in Fig. 2.

Fig. 2 CT scanning equipment showing control console, viewing screen and radiation proofed cabinet containing the X-ray source, turntable and detector screen. Courtesy Nikon Metrology UK Ltd.

The 3D image can be manipulated to provide slices through the object and these slices can be rotated so that the precise size and position of internal features can be accurately determined. Assembly or machining errors can be identified. Similarly, flaws within a welded joint or casting can be categorised and the size and position of inclusions, lack of fusion, cracks etc accurately measured, thus allowing an accurate Engineering Critical Assessment (ECA) to be carried out. The method can also be used to examine non-metallics – a couple of examples being the inspection of composite wind turbine blades for delaminations and the positioning of the cores in lost wax casting. The technique can also be used for accurate measurement of the internal and external surfaces of a component, the image being capable of being overlaid on a CAD drawing of the object. This is particularly useful where parts are provided by a number of suppliers to a manufacturer for subsequent assembly. The items can be checked against a CAD drawing before despatch and, if required, the manufacturer instantly provided with the results, thus reducing the risk of being supplied with non-fitting parts.  

Unfortunately, unlike real time radiography that has moved from being a non-portable method to construction site use, CT scanning requires the use of a protective enclosure containing the X-ray source, the detector screen and a precision turntable. The method is currently therefore confined to within a factory or laboratory environment.  

The final radiographic method is neutron radiography. Neutrons, like X- and gamma-radiation, will pass through solid materials and can be used to produce an image on photographic film or to react with a detector screen. Neutrons, however, unlike X- and gamma radiation, are strongly affected and absorbed by certain elements such as hydrogen, carbon and boron but to a far lesser extent by iron, nickel and aluminium. Neutron radiography is therefore useful in detecting the presence of material containing hydrogen or

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hydrocarbons – this includes corrosion products, water, explosives, oil and plastics. Composite items e.g. a combination of metal and plastic, can be non-destructively examined for assembly faults or manufacturing defects; explosive ordnance can be examined for complete filling and fluid flow analysis can be carried out in real time – oil flow in vehicle engines, fluid flow in piping systems.

Neutron radiography can be used with conventional film with real time techniques and for CT scanning. The neutrons are produced by particle accelerators or, with poorer image definition, from certain isotopes, primarily californium252. The method is currently not portable but development work is in progress to address this limitation.Ultrasonic Examination Part 1

Job Knowledge

Ultrasonic examination uses the same principles as the sonar used for the detection of submarines – a sound wave is emitted from a transmitter, bounces off any objects in its path and is reflected back to a receiver, somewhat similar to shining the beam of a torch at a mirror. Knowing the speed of sound in the material enables the distance of an object to be determined by measuring the time that elapses between the transmission of the sound pulse and detection of the “echo”. In welded components the examination is generally performed by moving a small probe containing both a transmitter and receiver over the item and displaying the echo on an oscilloscope screen. This is shown in Fig. 1 which illustrates a simple pulse-echo angle probe examination.

The oscillator sends pulses of electricity to a piezo-electric crystal, the pulse generator, embedded in the ultrasonic probe which causes it to vibrate at a very high (ultrasonic) frequency, well above any audible frequency and typically between 1Mega Hertz(MHz) and 15MHz. Ultrasonic probes used for weld examination have frequencies generally between 2MHz and 5Mhz, the lower frequency probes being used for the examination of coarse grained material or on rough surfaces, the higher frequency probes for the detection of fine defects such as cracks or lack of fusion. The ultrasonic vibrations are transmitted into the material to be tested using a “couplant” such as grease, paste or water which helps transmission of the vibrations. The better the surface finish then the better is the coupling and the more searching the examination – hence there is sometimes a requirement to grind smooth the weld cap and remove the root penetration bead on welded joints.

Once in the material the vibrations travel in a predictable path as a beam of sound pulses until they encounter an obstruction or interface such as a line of slag, porosity or a crack when most of the sound will be reflected - remember the analogy of the torch and mirror. Depending on the angle at which the beam strikes the obstruction some or all of the sound beam will be reflected back to the receiver in the probe. Here it vibrates a piezo-electric crystal; the electric signal that is generated is amplified, rectified and displayed on an oscilloscope screen.

The sound beam when it enters the object being scanned has a cross section approximately that of the transmitter but, like the beam of a torch, will diverge as shown in Fig. 1. As the beam travels through the material it also loses energy – it becomes attenuated. These effects need to be taken into account when the position and size of a defect is to be accurately determined.

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Fig 1. Schematic of Angle Probe Ultrasonic examination.

The oscilloscope screen in Fig. 1 shows on the vertical axis the signal height or amplitude and on the horizontal axis the time taken for the signal to return to the receiver and therefore distance from the transmitter. This method of examination is known as an “A scan” and is the most common method in use in industry for the examination of welded joints. In Fig 1 three signal peaks can be seen on the oscilloscope screen – one where the signal enters the sample, one reflected from the back face of the sample - the “back wall echo” - and, between the two, a reflection from some feature – a “reflector” such as a welding defect. The distance of this signal on the screen from the transmission pulse will give the distance of the reflector from the probe so a little simple geometry can be used to calculate the position and depth of the reflector within the block of material. Comparing the height of the signal with the signal from a known size of reflector enables the size of the feature to be determined.

There are two main types or modes of sound waves – longitudinal or compression waves which alternately compress and decompress the material in the direction of propagation and shear waves which vibrate the material at right angles to the direction of propagation. Which mode is produced depends upon the angle at which the sound beam enters the material. Probes that project the beam into the test piece at an angle normal (90degs) to the plate surface are known as compression probes and are ideally suited to the detection of defects such as plate laminations or for the measurement of plate/pipe thickness as shown in Fig 2.

Fig. 2. Compression Wave Probe

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To obtain the strongest reflected signal the beam should ideally strike the feature at 90O – flaws that lie parallel to the beam may be missed. This means that to examine a weld that may contain flaws laying in any number of orientations within the weld a range of different angle probes and scanning patterns must be used. To do this both compression and shear wave probes may be used, shear wave probes projecting the beam into the test piece at an angle, as shown in Fig. 1. Probes with angles of 30o, 45o, 60o and 70o are commercially available. Examples of standard probes are illustrated in Fig. 3.The angle of the probe is often selected to give the strongest signal from the defect of interest and for very high integrity welds all five probe angles may be used.

Fig 3. A 2.5MHz 70 degrees shear wave probe and a compression wave probe.

As shown in Fig. 4 the sound beam can be made to scan the full depth of a weld by moving the probe back and forth. At the half skip distance the beam would readily detect lack of side wall fusion along the left hand fusion line but may miss lack of side wall on the right hand fusion line. Moving the probe to the full skip distance so that the beam reflects off the back face enables the right hand fusion line to be scanned with the beam at the optimum angle to detect lack of side wall fusion.

Fig 4. Schematic of “skip distances”

To examine completely the weld there needs to be a number of such scanning patterns longitudinal and transverse to the weld, from both sides of the weld, from both plate surfaces and from half to full skip distance.  If all of these scans are carried out using all five probe types and two frequencies then it becomes

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a very lengthy and costly exercise! Such detailed examinations tend to be confined to items such as primary circuit nuclear components and highly critical offshore applications. Whilst many ultrasonic examinations are carried out with a manual operative moving the probe, viewing the results on the oscilloscope screen and manually recording the results the process can be mechanised with the probes mounted on a carriage and the results recorded electronically. This has become more prevalent as computing power has increased since the carriage may carry several probes and provides information on the carriage position and orientation. This data is then analysed and compared with an acceptance standard, enabling a weld to be sentenced automatically .

There are a number of advantages to ultrasonic testing:-

1. It is very good – and better than radiography - for the detection of planar defects such a lack of fusion and cracks

2. It can determine both the depth and position of defects.

3. It is readily portable and easy to use on site and in areas of restricted access.

4. Access is required to one side only.

5. There are none of the health and safety problems associated with radiography.

6. The result is immediately available.

But, as with any industrial process there are some disadvantages:-

1. Very skilled and conscientious operatives are required .2. The manual examination process is slow, laborious and tiring for the operative.

3. Surface breaking defects are difficult to detect.

4. Accurate sizing of small (<3mm) defects is difficult if not impossible.

5. The root region in a single sided full penetration weld is difficult to interpret.

6. The geometry of the joint can restrict the scanning pattern and impede accurate interpretation.

7. Interpretation is subjective and depends upon the operative’s skill and experience.

8. With manual scanning no permanent and objective record is produced.

The A-scan mentioned above is one method for reporting the results of the scan – there are in fact four methods identified as A-, B-, C- and D-scan. The A-scan method is the conventional way of presenting the results – signal amplitude vs distance; B-scan is a view looking along the length of the weld; C-scan is a plan view and D-scan a view from the side of the weld. These are illustrated in Fig.5

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Fig 5 Schematic of A-, B-, C- and D- scan results

Ultrasonic Examination Part 2

Job Knowledge

The previous article (127) explained the basic principles of ultrasonic examination. As to determine accurately the size and position of a feature it is necessary, with any measuring equipment, to calibrate the ultrasonic examination system. 

The type of calibration blocks (there are varying shapes and sizes to be used), depend on the application and the form and shape of the subject being tested. The calibration block should be made the same as the material being inspected and the artificially-induced flaw should closely resemble the actual flaw of concern. The best calibration block for calibrating ultrasonic testing equipment is one in the same grade of material and heat treatment condition as the production items and with a weld containing genuine flaws such as slag entrapment, porosity, lack of fusion, cracks etc. Techniques developed enable flaws of known sizes to be introduced into a welded joint. Such calibration blocks can be produced to validate the ultrasonic test method but are expensive and tend to be used only in applications such as nuclear vessel manufacture and critical offshore/process plant fabrication.

A number of standard calibration blocks are available with the shape and dimensions being specified in international standards such as ISO 2400, ISO 7963, ASME V and ASTM E164. Calibration of a compression wave probe used to measure thickness is simple and carried out using a stepped wedge calibration block. These calibration blocks have smooth, machined features and are not therefore truly representative of flaws in a welded component.

For calibrating equipment to be used to interrogate welded joints the calibration block needs to be more complex than a simple step wedge, with probably the two most common types illustrated in Fig. 1, the ISO

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2400 Number 1 block and the ISO 7963 Number 2 block. These are machined from steel to very closely controlled tolerances and contain a number of features that can be used to calibrate the ultrasonic equipment. The standard Number 1 block is 300mm long and 25 or 50mm thickness with a 100mm radius machined on one end. The test block also contains two drilled holes, 50 and 1.5mm in diameter and a flat bottomed machined notch. 

Fig 1: Number 1 and Number 2 calibration blocks

Smaller lighter blocks are useful for site use, and may be used to calibrate both compression and shear-wave probes for beam angle, time base, resolution and sensitivity. Sensitivity and resolution are terms frequently used – sensitivity is the ability to detect small flaws within the weld, resolution the ability to locate and separate individual flaws.

Weld discontinuity acceptance criteria are initially based on the height of the signal displayed on the oscilloscope screen. This is not as simple as it may appear since the ultrasonic beam is influenced by the microstructure of the metal through which it is propagating, becoming scattered and diffused - similar to car headlights in fog! As a general rule the larger the grain size the greater the scattering effect, the reflected beam becomes attenuated or decreased in strength the further away the reflector is from the ultrasonic probe. This must be taken into account when accepting or rejecting flaws within the weld – a 4MHz signal would lose some 0.02–0.03db per mm in steel. Fig 2 illustrates this decrease in amplitude or signal height with distance. 

Before calibrating the operator must select the frequency of the transducer as this determines the wavelength of the sound. The frequency has a significant effect on the ability to detect a flaw – a rule of thumb is that a flaw must be larger than one half the wavelength to be readily detectable.  

The ultrasonic operator will select a calibration block with some feature of known dimensions, often a 3mm diameter flat bottomed hole and the appropriate ultrasonic probe, these generally being specified in the relevant application code or standard. The height of the reflection at known distances from the probe would be determined and from this data would be drawn a distance amplitude correction (DAC) curve by joining the tips of the signals that can be seen in Fig 2. This provides a means of establishing a ‘reference level sensitivity’ as a function of distance from the ultrasonic probe and allows the signals from similar reflectors to be evaluated. 

The characteristics of an ultrasonic probe vary according to the size of the piezo-electric transducer and its frequency. It is therefore essential that each probe to be used to examine a welded component is individually calibrated and a DAC curve established for each different situation.

The contract specification, application code or acceptance standard specifies the relevant ultrasonic acceptance standard of height, length, position etc of the reflector. It is unwise to refer to a visual or radiographic acceptance standard in the absence of a relevant ultrasonic acceptance standard. An ultrasonic acceptance standard will state which reflectors are acceptable or unacceptable based on the amplitude of the signal compared with a DAC curve or other ultrasonic specific acceptance criteria. One such specification

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that refers to the DAC curve is ISO 11666 ‘NDT of welds – Ultrasonic testing – Acceptance levels’ which defines four levels:

the reference level ie the amplitude of the DAC curve at the relevant distance the evaluation level ie the amplitude at which the reflector must be examined more closely to

determine through thickness height and length of the discontinuity the recording level ie amplitude at which the size and position of the discontinuity must be recorded the acceptance level above which the discontinuity must be rejected – this may be above or below

the DAC curve. Any reflector with a signal below the evaluation level would be ignored

Fig. 2 The reduction in amplitude with distance

If, as the ultrasonic testing (U/T) probe is scanned across the surface of the component, and the amplitude of the signal exceeds the specified evaluation level, the U/T operator would need to investigate the reflector in detail to determine its size, orientation and position within the component. If the probe is moved transverse and parallel to the weld and rotated slightly, a skilled and experienced operator can often identify the flaw type – crack, lack of fusion, etc – by observing the changes in the shape of the pulse-echo on the oscilloscope screen.

To enable the operator to identify the position of a flaw it must be possible for the path and width of the beam to be visualised. Accurately dimensioned diagrams of the weld-cross section superimposed on what would be the beam path are required. This may be unnecessary in many situations but provides additional confidence in critical applications and may be a mandatory part of a written U/T procedure. 

The size of a reflector is generally determined by the ‘6db drop method’, as illustrated in Fig. 3.

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The operator moves the probe backwards and forwards at right angles to the axis of the reflector until the maximum amplitude response is found. This point is noted and the scanning continued until the amplitude of the signal has dropped by 6dB, this point also being recorded. From this the length or height of the reflector can be determined (Fig 3). If above the recording level this would be recorded on the U/T report before being compared with the acceptance standard for either acceptance or rejection.

It is impossible to measure accurately the size of a reflector using a manual scanning technique for a number of reasons. The speed of the sound within the component may vary due to changes in the microstructure and the cleanliness of the parent metal; the probe will be made to within dimensional tolerances, as will the calibration block and these will affect the accuracy of calibration; the beam width may vary; the couplant and surface condition of the component will affect the coupling and hence sound transmission; the surface of flaws within the weld are generally not flat, smooth reflective surfaces oriented at 90 degrees to the beam; the probe movement is measured manually with a rule or tape measure. The most important factors in achieving accurate, consistent and reproducible results are the skill, competence and integrity of the operator.

The accuracy of conventional manually-scanned pulse-echo ultrasonic examination carefully performed by a competent operator is around ±2mm. Such inaccuracy can be important when carrying out a fitness for service analysis, where the through thickness of a flaw is of critical importance. Some methods of achieving greater accuracy will be dealt with in the next article.Ultrasonics Examination: Part 3

Job Knowledge

The previous article   dealt with the manual scanning method of ultrasonic examination stating that accurate determination of weld flaw size and position - to within ±2mm -  was difficult, if not impossible, in most circumstances. Methods developed now enable flaw sizes to be determined with accuracy better than ±1mm. This article will look at two of these techniques; time of flight diffraction (TOFD) and phased array ultrasonic testing (PAUT). 

Conventional manually scanned ultrasonic testing normally uses a single fixed angle and frequency probe;

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the position and size of a flaw being determined by the amplitude of the signal reflected from the flaw and presented on an oscilloscope screen (Fig 3 in Job Knowledge 128). This is a somewhat unreliable method as the amplitude of the signal and therefore an estimate of its size depends on the orientation of the flaw. TOFD uses two angled compression wave probes mounted on a frame so that they are a fixed distance apart; one a transmitter, the other a receiver. The probes are positioned either side of a weld as shown in Fig 1. In a flaw-free weld two sound waves will be detected by the receiver – one that travels along the surface of the weld, the lateral wave, and one reflected from the back wall. When a flaw is present (for example a crack as shown in Fig 1) the pulse emitted by the transmitter is diffracted or scattered from the tip of a flaw and this diffracted signal is picked up by the receiver. The time of flight of the signal is measured and compared with that of the lateral wave, a simple calculation enables the position of the tip of the flaw to be determined. Moving the probes in a predetermined scanning pattern enables the other end of the flaw to be detected so the flaw size can be established in both the trough thickness and longitudinal directions. The calculations are performed automatically by the software program within the equipment and the flaws displayed as a black and white A scan image. 

Figure 1: TOFD

Various scanning patterns may be used so that the results can be presented as A-scan, B-scan, looking along the weld length or D-scan, a side view. (For a description of A, B, C and D scans see Job Knowledge No. 127).

TOFD is regarded as the best method for the detection and sizing of planar, through-thickness flaws. One limitation is the detection of small surface breaking flaws on the scanning surface as these tend to be lost in the lateral wave response, although this may be not too significant as most surface breaking defects can be readily detected using MPI or liquid penetrant methods. 

The rapid progress of electronics and computing power has enabled complex methods of scanning and data processing to be developed. This has culminated in phased array ultrasonic testing (PAUT) which, as the name suggests, uses an array of small transducers unlike the conventional manual A-scan probe with only one transducer. A single PAUT probe may contain between several tens and several hundreds of transducers. These small transducers are computer controlled and can be pulsed independently in a set sequence or phase; the pulses of sound interfering with each other to produce a sound beam of a certain angle, see Fig. 2. By varying the time and pattern of the pulses, the angle and shape of the beam can be varied so that the beam can be steered electronically, sectorial scanning or S-scan. 

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Figure 2: Illustration of the sector scan composed of many A-scans from the beams being steered through a range of angles. Note that, in addition to steering the beam, the focal law may also be focusing the sound

field to improve defect detection and resolution

The benefits of this technology compared with conventional single transducer scanning are that the beam can be steered and focused with a single probe. Beam steering enables the beam to be swept through an object without moving the probe, the reflected data being processed to provide a visual image of a cone shaped slice through the object. Moving the probe enables a large number of slices to be assembled to provide a three dimensional image – a good example is the use in medical diagnostics to examine the functioning of the heart in real time. 

For the non-destructive examination of welds this ability to inspect a weld with multiple angle beams from a single probe means that the probability of detecting flaws is greatly increased. It is also possible to focus the beam electronically at multiple depths to improve the ability to accurately determine the size and position of weld flaws. 

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Figure 3: Phased array results showing A-scan, B-scan and S-scan of a nozzle to shell weld

The small probe size and the ability to manipulate the beam without moving the probe enables inspections in limited access or of components of a complex shape. Cost is also a factor – although the probes and the processing/display units are more expensive than the single transducer equipment, the time to perform a scan can be substantially reduced. Work carried out by TWI suggests that a phased array scan can take 20% of the time for a conventional scan with better coverage although the off-line interpretation of the results may take longer. 

The results may be presented as S-, A-, B- or C-scans, enabling better interpretation. The results of a phased array examination of a single sided nozzle to shell weld is given in Fig. 3. The weld shape is given by the red lines superimposed on the S-scan display. This is a single sided weld, the lower half of the image being a mirror of the weld. Whilst the scanning operation may be performed automatically by mechanised manipulating equipment and the accuracy may be better than ±1mm, the interpretation must be carried out by experienced and skilled personnel trained specifically in the interpretation of phased array scanning results. An investigation by TWI showed that the skill of the individual carrying out the interpretation was by far the most important factor in producing reproducible and accurate results.   

Scanning can be performed manually or with the probe attached to a carriage. A typical application is the examination of pipe butt welds using orbiting crawler tractors 

Such dedicated and robust mechanised equipment is readily available for site use, replacing radiography and giving benefits in terms of cost, flaw detection and health and safety issues.  

Oxyfuel cutting - process and fuel gases

Job Knowledge

Mechanised oxyacetylene cutting system

The oxyfuel process is the most widely applied industrial thermal cutting process because it can cut thicknesses from 0.5mm to 250mm, the equipment is low cost and can be used manually or mechanised. There are several fuel gas and nozzle design options that can significantly enhance performance in terms of cut quality and cutting speed.

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Process fundamentals

The cutting process is illustrated in Fig. 1. Basically, a mixture of oxygen and the fuel gas is used to preheat the metal to its 'ignition' temperature which, for steel, is 700°C - 900°C (bright red heat) but well below its melting point. A jet of pure oxygen is then directed into the preheated area instigating a vigorous exothermic chemical reaction between the oxygen and the metal to form iron oxide or slag. The oxygen jet blows away the slag enabling the jet to pierce through the material and continue to cut through the material.

Fig.1. Diagram of oxyacetylene cutting process

There are four basic requirements for oxy-fuel cutting:

the ignition temperature of the material must be lower than its melting point otherwise the material would melt and flow away before cutting could take place

the oxide melting point must be lower than that of the surrounding material so that it can be mechanically blown away by the oxygen jet

the oxidation reaction between the oxygen jet and the metal must be sufficient to maintain the ignition temperature

a minimum of gaseous reaction products should be produced so as not to dilute the cutting oxygenAs stainless steel, cast iron and non-ferrous metals form refractory oxides ie the oxide melting point is higher than the material, powder must be injected into the flame to form a low melting point, fluid slag.

Purity of oxygen

The cutting speed and cut edge quality are primarily determined by the purity of the oxygen stream. Thus, nozzle design plays a significant role in protecting the oxygen stream from air entrainment.

The purity of oxygen should be at least 99.5%. A decrease in purity of 1% will typically reduce the cutting speed by 25% and increase the gas consumption by 25%.

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Choice of fuel gas

Fuel gas combustion occurs in two distinct zones. In the inner cone or primary flame, the fuel gas combines with oxygen to form carbon monoxide and hydrogen which for acetylene, the reaction is given by

2C2H2 + 2O2 → 4CO + 2H2

Combustion also continues in the secondary or outer zone of the flame with oxygen being supplied from the air.

4CO+2H2 +3O2 → 4CO2 +2H2O

Thus, fuel gases are characterised by their

flame temperature - the hottest part of the flame is at the tip of the primary flame (inner cone) fuel gas to oxygen ratio - the amount of fuel gas required for combustion but this will vary according

to whether the flame is neutral, oxidising or reducing heat of combustion - heat of combustion is greater in the outer part of the flameThe five most commonly used fuel gases are acetylene, propane, MAPP (methylacetylene-propadiene), propylene and natural gas. The properties of the gases are given in the Table. The relative performance of the fuel gases in terms of pierce time, cutting speed and cut edge quality, is determined by the flame temperature and heat distribution within the inner and out flame cones.

Acetylene

Acetylene produces the highest flame temperature of all the fuel gases. The maximum flame temperature for acetylene (in oxygen) is approximately 3,160°C compared with a maximum temperature of 2,828°C with propane. The hotter flame produces more rapid piercing of the materials with the pierce time being typically one third that produced with propane.

The higher flame speed (7.4m/s compared with 3.3m/s for propane) and the higher calorific value of the primary flame (inner cone) (18,890kJ/m3 compared with 10,433 kJ/m3 for propane) produce a more intense flame at the surface of the metal reducing the width of the Heat Affected Zone (HAZ) and the degree of distortion.

Propane

Propane produces a lower flame temperature than acetylene (the maximum flame temperature in oxygen is 2,828°C compared with 3,160°C for acetylene). It has a greater total heat of combustion than acetylene but the heat is generated mostly in the outer cone (see Table). The characteristic appearance of the flames for acetylene and propane are shown in Figs.2 and 3 where the propane flame appears to be less focused. Consequently, piercing is much slower but as the burning and slag formation are effected by the oxygen jet, cutting speeds are about the same as for acetylene.

Propane has a greater stoichiometric oxygen requirement than acetylene; for the maximum flame temperature in oxygen, the ratio of the volume of oxygen to fuel gas are 1.2 to 1 for acetylene and 4.3 to 1 for propane.

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Fig.2. Ocyacetylene gas jet and nozzle design

Fig.3. Propane gas jet and nozzle design

MAPP

MAPP gas is a mixture of various hydrocarbons, principally, methylacetylene and propadiene. It produces a relatively hot flame (2,976°C) with a high heat release in the primary flame (inner cone) (15,445kJ/m3), less than for acetylene (18,890kJm3) but much higher than for propane (10,433kJm3). The secondary flame (outer cone) also gives off a high heat release, similar to propane and natural gas. The combination of a lower flame temperature, more distributed heat source and larger gas flows compared with acetylene results in a substantially slower pierce time.

As MAPP gas can be used at a higher pressure than acetylene, it can be used for underwater cutting in deep water as it is less likely to dissociate into its components of carbon and hydrogen which are explosive.

Propylene

Propylene is a liquid petroleum gas (LPG) product and has a similar flame temperature to MAPP (2896°C compared to 2,976°C for MAPP); it is hotter than propane, but not as hot as acetylene. It gives off a high heat release in the outer cone (72,000kJ/m3) but, like propane, it has the disadvantage of having a high stoichiometric fuel gas requirement (oxygen to fuel gas ratio of approximately 3.7 to 1 by volume).

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Natural Gas

Natural gas has the lowest flame temperature similar to propane and the lowest total heat value of the commonly used fuel gases, eg for the inner flame 1,490kJ/m3 compared with 18,890kJ/m3 for acetylene. Consequently, natural gas is the slowest for piercing.

Table : Fuel Gas Characteristics

Fuel GasMaximum FlameTemperature °C

Oxygen to fuel gasRatio (vol)

Heat distributionkJ/m3

      Primary Secondary

Acetylene 3,160 1.2:1 18,890 35,882

Propane 2,828 4.3:1 10,433 85,325

MAPP 2,976 3.3:1 15,445 56,431

Propylene 2,896 3.7:1 16,000 72,000

Hydrogen 2,856 0.42:1 - -

Natural Gas 2,770 1.8:1 1,490 35,770

Cutting processes - application of oxyfuel cutting

Job Knowledge

Rough-cut gear wheel cut by oxyacetylene

Oxyfuel is one of the most widely used cutting processes with the following benefits:

Low cost equipment Basic equipment suitable for cutting, gouging and other jobs such as welding and heating Portable, suitable for site work Manual and mechanised operations

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Mild and low alloy steels (but not aluminium or stainless steel) Wide range of thickness (typically from 1mm to 1000mm)It is therefore not surprising that the process can be used for a diverse range of applications from manual rough severing and scrap cutting to precision contour cutting in fully automated systems. Here, the process application is described including the choice of fuel gas and nozzle design to maximise performance. Best practice to ensure adequate quality of the cut surface is also included.

Choice of fuel gas

Basically, a mixture of oxygen and a fuel gas (acetylene, propane, MAPP propylene or methane) is used to preheat the metal to its 'ignition' temperature which is well below its melting point. A jet of pure oxygen is then directed into the preheated area which burns through the spot and the resulting molten metal and slag are removed by the high velocity oxygen stream. The cutting speed is primarily determined by the oxygen jet but as the outer fuel gas/oxygen flame determines the rate of preheating, the choice of fuel gas has a significant influence on the time taken to initiate the cutting operation. This is especially important if the designed cut begins by piercing.

The choice of fuel gas is largely made on cost, performance, ease of use and whether it is a manual or mechanised operation. However, in making the choice it should be noted that in a typical application the cost is made up of approximately:

50% overheads 30% handling labour 18% cutting labour 1-2% gasConsideration should, therefore, be given to the choice of fuel gas type and nozzle design to speed up the initiation of the cutting operation. Labour costs can be reduced by decreasing the pierce time and/or increasing the cutting speed. Typical flame temperatures and fuel gas to oxygen ratios are shown in Fig. 1. Generally, fuel gases which generate a higher flame temperature and require a lower oxygen to fuel gas ratio, will speed up the cutting operation.

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Fig. 1. Flame temperature and the fuel gas to oxygen ratio

Acetylene

Acetylene produces the highest flame temperature of all the fuel gases and generates a highly focused flame. As the pierce time is approximately one third that achieved with propane, it should be used when the pierce time is a significant proportion of the total cutting time, for example, short cuts and multi-pierce cutting operations.

The high temperature (maximum flame temperature in oxygen is 3160°C), highly focused flame makes the oxyacetylene process ideal for cutting thin sheets with minimum distortion and for bevel cutting. However, the high cost and low heat generation make it less suitable for general heating of large plates.

Propane

Propane is low cost and has the advantage of being available in bulk supplies. The flame temperature is lower than for acetylene (the maximum flame temperature in oxygen is 2828°C compared with 3160°C for acetylene) which makes piercing much slower. However, it can tolerate a greater nozzle to workpiece distance which reduces the risk of molten metal splashing back onto the nozzle and causing a 'backfire'.

For similar nozzle designs, cutting speeds for oxypropane and oxyacetylene are similar. Advantages claimed for propane are smooth cut edge, less slag adhesion and lower plate edge hardening because of the lower flame temperature. The heat affected zone is much wider than for oxyacetylene.

MAPP

MAPP gas, which is a mixture of various hydrocarbons, principally, methylacetylene and propadiene, produces a relatively hot flame (2976°C). However, the lower calorific value of the inner cone compared with acetylene gives a slightly slower pierce time.

The gas is seen as an alternative to acetylene with greater tolerance to torch distance variation because of the more uniformly distributed heat between the inner and the outer cones.

Only acetylene, hydrogen and MAPP have sufficiently high flame temperature for underwater cutting. But as acetylene has a limited outlet pressure, MAPP is the only gas other than hydrogen that can be used for cutting in deep water.

Propylene

Propylene is a liquid petroleum gas (LPG) product and has a similar flame temperature to MAPP (2896°C compared to 2976°C for MAPP). It gives off a high heat release in the outer cone (72,000 kJ/m3) but, like propane, it has the disadvantage of having a high stoichiometric oxygen requirement (oxygen to fuel gas ratio of approximately 3.7 to 1 by volume).

Methane

Methane has the lowest flame temperature similar to propane and the lowest total heat value of the commonly used fuel gases. Consequently, natural gas is the slowest for piercing.

Cutting torch

The cutting torch design can be either nozzle mix or injector. In the nozzle mix torch, the fuel gas and pre-heat oxygen are mixed in the nozzle. In the injector torch, the pre-heat gases mix either in the body of the torch, within the gas delivery tubes, or within the head of the torch. Injector torches have the advantage of

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being able to use the higher pressure of oxygen to pull the fuel into the torch. This allows the torch to be used at low fuel gas pressures or with large pressure drops such as those experienced through long hose lengths.

Nozzle

The primary functions of the nozzle are to provide:

a method of preheating the metal to its ignition temperature a jet of oxygen to react with the material to be cut and at a flow rate sufficient to blow away the slagEach torch should be fitted with the appropriate nozzle for the type of fuel gas. Nozzles can be of a one- or two-piece design. The nozzle type will depend on:

fuel gas manual or machine operation manufacturer's preference Acetylene nozzles are usually one-piece but two-piece nozzles similar to those for other fuel gases are produced for machine cutting.

The diameter of the cutting oxygen hole is selected according to the material thickness. There are two types of nozzle; standard and high speed. The standard nozzle usually has a parallel sided, central bore for the oxygen jet, which is surrounded by an annulus or a ring of smaller diameter ports for the pre-heating gas mixture, Fig. 2. There are many designs and arrangements of the preheating ports that focus the flame for heating and to protect the oxygen jet from air entrainment.

Fig. 2. Standard nozzle with central bore for oxygen jet and a ring of ports for the pre-heating gas mixture

High-speed nozzles are capable of being used with higher oxygen pressures, up to 10 bar. The essential difference is that the cutting oxygen is forced through a convergent / divergent orifice which speeds up the gas flow rate to near supersonic levels. High-speed nozzles are primarily used in mechanised equipment to exploit the higher speeds for cutting long lengths.

Best practice

Cutting conditions are normally set to produce an acceptable cut surface finish for the application but at the highest cutting speed. It is, therefore, essential that consideration is given to the following settings for the material thickness and the cutting speed:

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nozzle distance - too high or too low will disturb oxygen flow preheat flame - too high a flow can cause top edge melting cutting oxygen - too low a flow can cause poor slag removal - too high a flow can result in poor cut finish

The typical appearances of a good and poor quality cut surface for manual cutting are shown in Fig.3. The principal features are described together with their cause and remedial measures necessary to produce the ideal square edge, smooth surface cut.

Ideal Cut Profile

Features 

Square edge, smooth cut surface, underside free of slag, small drag lines

Cutting Too Fast

Features 

Coarse drag lines at angle to surface with excessive amount of slag sticking to bottom edge of plate

Cause Oxygen jet trailing with insufficient oxygen reaching bottom of the cut

Too high nozzle to plate distance

Features 

Uneven cut surface with heavy melting of top edge, coarse drag lines at bottom cut surface

Cause Preheat is not focused on plate surface, oxygen jet easily disturbed

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Too High Oxygen Flow

Features 

Excessive slag adhering to cut face, local gouging, excessive top edge melting

Cause Turbulence between the preheat flame and the cutting jet

Fig. 3. Best practice guide for hand cutting

Cutting processes - plasma arc cutting - process and equipment considerations

Job Knowledge

Photo courtesy: Goodwin Plasma

The plasma arc process has always been seen as an alternative to the oxy-fuel process. In this part of the series the process fundamentals are described with emphasis being placed on the operating features and the advantages of the many process variants.

Process fundamentals

The plasma arc cutting process is illustrated in Fig. 1. The basic principle is that the arc formed between the electrode and the workpiece is constricted by a fine bore, copper nozzle. This increases the temperature and velocity of the plasma emanating from the nozzle. The temperature of the plasma is in excess of 20 000°C and the velocity can approach the speed of sound. When used for cutting, the plasma gas flow is increased so that the deeply penetrating plasma jet cuts through the material and molten material is removed in the efflux plasma.

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Fig.1. The plasma arc cutting process

The process differs from the oxy-fuel process in that the plasma process operates by using the arc to melt the metal whereas in the oxy-fuel process, the oxygen oxidises the metal and the heat from the exothermic reaction melts the metal. Thus, unlike the oxy-fuel process, the plasma process can be applied to cutting metals which form refractory oxides such as stainless steel, aluminium, cast iron and non-ferrous alloys.

Power source

The power source required for the plasma arc process must have a drooping characteristic and a high voltage. Although the operating voltage to sustain the plasma is typically 50 to 60V, the open circuit voltage needed to initiate the arc can be up to 400V DC.

On initiation, the pilot arc is formed within the body of the torch between the electrode and the nozzle. For cutting, the arc must be transferred to the workpiece in the so-called 'transferred' arc mode. The electrode has a negative polarity and the workpiece a positive polarity so that the majority of the arc energy (approximately two thirds) is used for cutting.

Gas composition

In the conventional system using a tungsten electrode, the plasma is inert, formed using either argon, argon-H2 or nitrogen. However, as described in Process variants, oxidising gases, such as air or oxygen, can be used but the electrode must be copper with hafnium.

The plasma gas flow is critical and must be set according to the current level and the nozzle bore diameter. If the gas flow is too low for the current level, or the current level too high for the nozzle bore diameter, the arc will break down forming two arcs in series, electrode to nozzle and nozzle to workpiece. The effect of 'double arcing' is usually catastrophic with the nozzle melting.

Cut quality

The quality of the plasma cut edge is similar to that achieved with the oxy-fuel process. However, as the plasma process cuts by melting, a characteristic feature is the greater degree of melting towards the top of the metal resulting in top edge rounding, poor edge squareness or a bevel on the cut edge. As these

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limitations are associated with the degree of constriction of the arc, several torch designs are available to improve arc constriction to produce more uniform heating at the top and bottom of the cut.

Process variants

The process variants, Figs. 2a to 2e, have principally been designed to improve cut quality and arc stability, reduce the noise and fume or to increase cutting speed.

Dual gas

Fig.2a. dual gas

The process operates basically in the same manner as the conventional system but a secondary gas shield is introduced around the nozzle, Fig. 2a. The beneficial effects of the secondary gas are increased arc constriction and more effective 'blowing away' of the dross. The plasma forming gas is normally argon, argon-H2 or nitrogen and the secondary gas is selected according to the metal being cut.

Steel

air, oxygen, nitrogen

Stainless steel

nitrogen, argon-H2, CO2

Aluminium

argon-H2, nitrogen / CO2

The advantages compared with conventional plasma are:

Reduced risk of 'double arcing' Higher cutting speeds Reduction in top edge rounding

Water injection

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Fig.2b. water injection

Nitrogen is normally used as the plasma gas. Water is injected radially into the plasma arc, Fig. 2b, to induce a greater degree of constriction. The temperature is also considerably increased, to as high as 30,000°C.

The advantages compared with conventional plasma are:

Improvement in cut quality and squareness of cut Increased cutting speeds Less risk of 'double arcing' Reduction in nozzle erosion

Water shroud

Fig.2c. water shrouded

The plasma can be operated either with a water shroud, Fig. 2c, or even with the workpiece submerged some 50 to 75mm below the surface of the water. Compared with conventional plasma, the water acts as a barrier to provide the following advantages:

Fume reduction Reduction in noise levels Improved nozzle lifeIn a typical example of noise levels at high current levels of 115dB for conventional plasma, a water shroud was effective in reducing the noise level to about 96dB and cutting under water down to 52 to 85dB.

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As the water shroud does not increase the degree of constriction, squareness of the cut edge and the cutting speed are not noticeably improved.

Air plasma

Fig.2d. air plasma

The inert or unreactive plasma forming gas (argon or nitrogen) can be replaced with air but this requires a special electrode of hafnium or zirconium mounted in a copper holder, Fig. 2d. The air can also replace water for cooling the torch. The advantage of an air plasma torch is that it uses air instead of expensive gases.

It should be noted that although the electrode and nozzle are the only consumables, hafnium tipped electrodes can be expensive compared with tungsten electrodes.

High tolerance plasma

Fig.2e. high tolerance

In an attempt to improve cut quality and to compete with the superior cut quality of laser systems, High Tolerance Plasma Arc cutting (HTPAC) systems are available which operate with a highly constricted plasma. Focusing of the plasma is effected by forcing the oxygen generated plasma to swirl as it enters the plasma orifice and a secondary flow of gas is injected downstream of the plasma nozzle, Fig. 2e. Some systems have a separate magnetic field surrounding the arc. This stabilises the plasma jet by maintaining the rotation induced by the swirling gas. The advantages of HTPAC systems are:

Cut quality lies between a conventional plasma arc cut and laser beam cut Narrow kerf width

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Less distortion due to smaller heat affected zone HTPAC is a mechanised technique requiring precision, high-speed equipment. The main disadvantages are that the maximum thickness is limited to about 6mm and the cutting speed is generally lower than conventional plasma processes and approximately 60 to 80% the speed of laser cutting. 

Cutting processes - laser cutting

Job Knowledge

Cut section of ellipse in flat plate

Coined from the words Light Amplification by Stimulated Emission ofRadiation lasers have been a byword for efficiency and quality in materials processing since their advent in the sixties.

They offered an entirely new form of energy which in turn lent itself to uses in manufacturing, medicine and communications. Able to heat, melt and even vaporise material lasers are seen as the ideal medium for combining intense but controllable energy.

By far the most popular use of the laser, particularly the carbon dioxide laser, is for cutting.

Laser cutting

It is largely a thermal process in which a focused laser beam is used to melt material in a localised area. A co-axial gas jet is used to eject the molten material from the cut and leave a clean edge.

A continuous cut is produced by moving the laser beam or workpiece under CNC control.

The process also lends itself to automation with offline CAD/CAM systems controlling either 3-axis flat bed systems or 6-axis robots for three dimensional laser cutting.

The improvements in accuracy, edge squareness and heat input control means that other profiling techniques such as plasma cutting and oxy-fuel cutting are being replaced by laser cutting.

Cutting characteristics Benefits

Cuts carbon manganese steels up to 20mm Cuts stainless steel up to 12mm Cuts aluminium up to 10mm Cuts brass and titanium

High quality cut - no finishing Ultra flexible - simple or complex parts Non contact - no surface blemishing Quick set up - small batches

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Cuts thermoplastics, wood and many non-metals Low heat input - small HAZ, low distortion Lends itself to nearly all materials

What's the relationship between the lens used and the thickness of cut?

The laser cutting process involves focusing a laser beam, usually with a lens, to a small spot which has sufficient power density to produce a laser cut.

The lens is defined by its focal length, which is the distance from the lens to the focused spot. However, the critical factors which determine the selection of the lens are the focused spot diameter, d, and the depth of focus, L.

The depth of focus is the effective distance over which satisfactory cutting can be achieved. It can be defined as the distance over which the focused spot size does not increase beyond 5%.

For a given beam diameter, as the focal length becomes shorter the focused spot diameter and the depth of focus also both become smaller. The size of the actual spot is also dependent on the raw beam diameter, D. As this increases, for a given lens, the focused spot size decreases.

To allow comparison between lasers with different beam diameters we therefore use a factor called the focus f-number, which is the focal length, F, divided by the incoming raw beam diameter, D.

As we are generally unable to alter the raw beam diameter we select the correct lens to give us a focus beam of the required type.

The requirements for cutting are high power density, and therefore small focused spot size but with a long depth of focus, and therefore the ability to process thicker materials with a reasonable tolerance to focus position variation.

These two requirements are in conflict with each other and therefore a compromise must be made. The only other consideration is that the shorter the focal length, the closer the lens is to the workpiece, and therefore more likely to be damaged by spatter from the cutting process.

For typical CO2 laser cutting systems focal lengths can be selected in the range from 21/2 inches up to 10 inches, which are equivalent to f-numbers between two and ten, depending on the beam diameter.

In practice a 5 inch lens could cut up to around half inch thick steel before a longer lens would be required to provide a greater depth of focus. However on thin sheet material, for example 1mm, a shorter focal length may offer significantly higher cutting speeds, or allow more intricate detail to be produced.

In fact it would be possible to optimise focal length for each material thickness, but this would involve additional set up time when changing from one job to the next, which would have to be balanced against the increased speed. In reality lens changing is avoided and a compromise cutting speed used, unless a specific job has special requirements .

Just how flexible are they?

Most laser cutting machines are 3-axis systems, that is X-Y, two dimensional positioning control with a Z-axis height control.

There are however a number of ways of achieving the X-Y movement, either moving the laser head, moving the workpiece or a combination of both.

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The most popular approach is known as a 'flying optics' system where the workpiece remains stationary and mirrors are moved in both X and Y axes. The advantages of this approach are that the motors are always moving a known, fixed mass. This can often be much heavier than the workpiece, but it is easier to predict and control.

As the workpiece is not moved, this also means that there is no real limit to sheet weight. The disadvantage of flying optics is the variation in beam size, as a laser beam is never perfectly parallel, but actually diverges slightly as it leaves the laser.

This means that without controlling the divergence, there may be some variation in cutting performance between different parts of the table, due to a change in raw beam size. This effect can be reduced by adding a re-collimating optic, or some systems even use adaptive mirror control.

The alternative is a 'fixed optic' system where the laser head remains stationary and the workpiece is moved in both X and Y axes. This is the ideal situation optically, but the worst situation mechanically, especially for heavier sheets.

For relatively light sheet weights, a fixed optic system can be a viable option, but as the sheet weight increases, accurately positioning the material at high speed can be a problem.

The third option is known as a 'hybrid' system, where the laser head is moved in one axis and the material moved in the other axis. This is often an improvement over fixed optics, but still suffers from difficulties with heavier sheet weights.

What difficulties does reflection cause?

Amada LCV laser cutting machine with autostorage and pallet changer system Courtesy of Amada UK Ltd

All metals are reflective to CO2 laser beams, until a certain power density threshold value is reached.

Aluminium is more reflective than carbon manganese steel or stainless steel and has the potential to cause damage to the laser itself.

Most laser cutting machines use a laser beam aligned normal to a flat sheet of material. This means that should the laser beam be reflected by the flat sheet it can be transmitted back through the beam delivery optics, and into the laser itself, potentially causing significant damage.

This reflection does not come entirely from the sheet surface, but is caused by the formation of a molten pool which can be highly reflective. For this reason simply spraying the sheet surface with a non-reflective coating will not entirely eliminate the problem.

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As a general rule the addition of alloying elements reduces the reflectivity of aluminium to the laser, so pure aluminium is harder to process than a more traditional 5000 series alloy.

With good, consistent cutting parameters the likelihood of a reflection can be reduced to almost zero, depending on the materials used. However it is still necessary to be able to prevent damage to the laser while developing the conditions or if something goes wrong with the equipment.

The 'aluminium cutting system' which most modern equipment uses is actually a way of protecting the laser rather than an innovative technique for cutting. This system usually takes the form of a back reflection system that can detect if too much laser radiation is being reflected back through the optics.

This will often automatically stop the laser, before any major damage is caused. Without this system there are risks with processing aluminium as there is no way of detecting if potentially hazardous reflections are occurring.

Welding costs

The previous Connect article, number 95, dealt with the methods of determining the weight of deposited weld metal in a joint, enabling the cost of welding consumables to be calculated.

This is obviously the first step towards calculating the cost of actually making a welded joint but there are many other factors that need to be considered but which are beyond the scope of these articles.

The most significant of these costs is the overhead; the cost of providing a welding workshop or site and the costs of managing and running the organisation.

These costs are dependent on the accounting practices of the organisation. They comprise factors such as rent, rates, bank interest, cost of indirect workers, iethose not directly involved in fabricating, depreciation of plant etc. In addition, other accounting decisions (for example, where the costs of machining and assembly are absorbed) may affect the decisions on which is the most cost-effective joining method.

One of the most significant costs is that of labour and this inevitably varies with industry, time and country. The costs mentioned above cannot generally be influenced by the decisions made by a welding engineer. These articles will therefore concentrate on those aspects of welding activities that are not subject to accounting practices, overhead or labour costs.

There are many costs, other than the cost of depositing weld metal that will affect the price of a welded fabrication.

The work done by the designer in designing the most cost-effective joint in an item that can be placed in the most advantageous position for welding will have major effect on costs. For example, the type of joint preparation the designer selects; a single or double-V preparation can be flame cut, a J-preparation must be machined and is generally far more expensive. A machined J-preparation, however, may have less volume than a single-V, depending on thickness; will be more accurate and therefore quicker to assemble within tolerance and may result in a lower repair rate leading to a lower cost than the V preparation.

Costs that are directly affected by welding engineering decisions, in addition to the cost of actually depositing weld metal, are therefore; joint preparation, assembly time (which includes positioning in any jig or fixture and tacking), cleaning and dressing the weld, removal from jigs or fixture, post weld heat treatment, costs of non-destructive testing and cost of repairs.

The amount of weld metal deposited is rarely the same as the amount of filler metal purchased. This is the result of losses when, for example, GMAW or submerged arc welding wire is trimmed back to the contact

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tip, when the wire reel runs out and the length of wire between the drive roll and the contact tip is scrapped or the wire or reel is damaged.

Such losses tend to be quite small but this is not the case with coated electrodes. Damaged flux coatings, incorrectly stored electrodes and the stub ends discarded by the welder all contribute to as much as a third of the purchased weight of manual metallic arc electrodes being scrapped. Some electrode manufacturers' catalogues give figures for these losses which can vary depending on electrode type and diameter.

To assist in calculating the amount of welding consumables to be purchased Table 1 gives some multiplication factors for the more common arc welding processes. The weight of weld metal in the joint should be multiplied by this factor to give the amount of welding consumable required. These figures assume good housekeeping and shop floor discipline such that consumables are not wasted or scrapped unnecessarily.

Table 1 Multiplication factor. Weight of weld metal to give the weight of filler metal required.

Arc welding process Multiplication factorMMA (SMAW) 1.5TIG (GTAW) 1.1MIG/MAG (GMAW) 1.05Sub Arc (SAW) 1.02FCAW 1.2MCAW 1.1The other consumables in this cost equation are shielding gases or flux.

The conventional shoulder height welding gas cylinder contains approximately 10,000 litres of shielding gas at a pressure of 200bar. As the gas flow rates normally used in production are around 12 to 15 litres per minute, this typical cylinder should provide in the region of 10 to 12 hours of welding time, allowing for losses at the beginning and end of the arcing period.

The rate of flux consumption in submerged arc welding is approximately 1kg of flux for every 1kg of deposited weld metal. This assumes good housekeeping and an efficient flux recirculation system. Calculation of the amount required (and hence the cost) of these consumables is therefore relatively straightforward.

The cost of the welder's time to weld a joint does not depend solely on the deposition rate of the process. A most important factor in determining the time required by the welder is what is known as the 'duty cycle' or 'operating factor'. This is a percentage figure giving the amount of time that the arc is burning and weld metal is being deposited versus the total time that the welder is working.

Table 2 gives some figures for the more common arc welding processes. Note that these do NOT include set-up or assembly time and individual circumstances can increase or decrease these figures.

Table 2 Duty cycles for arc welding processes

Arc welding process Duty cycle %MMA (SMAW) 15 - 30TIG (GTAW) 25 - 40Mechanised TIG 80 - 90MIG/MAG (GMAW) 30 - 45

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Mechanised MIG/MAG 80 - 90Sub Arc (SAW) 80 - 95FCAW 25 - 45Mechanised FCAW 70 - 85MCAW 30 - 45The lost time in this figure can be accounted for by considering all of the other activities that the welder performs. In MMA welding, for example, time is required for tacking, de-slagging and cleaning a weld pass, for changing electrodes, for changing position, for rest breaks and for removal of the item from a fixture. Similar activities need to be performed using the other welding processes.

Increasing the duty cycle is therefore one method of increasing productivity, either by organising the shop floor such that lost time is reduced or by the use of a higher duty cycle process. However, remember that the arcing time may well be only a very small proportion of the total time to manufacture and attention to other aspects of the manufacturing cycle may give better returns than simply increasing the welding duty cycle.

Reference to Table 2 also suggests that mechanisation is one method to increase the duty cycle. Caution needs to be exercised, however, if the total (floor to floor) time is to be reduced. For one-off or small batch items the time taken to prepare and set up a mechanised system to weld the item may be longer than that taken to weld using a manual process. Note also that if a mechanised system is used, the duty cycle may in fact decrease, as the welding speed is increased and the weld is completed in a shorter time although the number of items welded per day will increase. It is therefore essential to consider the complete manufacturing cycle to achieve the most cost effective solution.

Welding costs - continued

Flux cored arc welding

The previous two Connect articles dealt with the mechanics of costing a weld: how to calculate the weld volume and how to calculate the amount of welding consumables required to fill a weld preparation.

The final step in costing a weld is to determine the length of time to deposit this weight of weld metal. This is obviously a function of the deposition rate of the process. The deposition rate is generally expressed as kgs/hr or lbs/hr deposited at a given welding current, welding continuously and without any breaks for electrode changing or deslagging.

The deposition rate will be affected by many factors and it will not be possible within the limitations of these articles to list the precise deposition rates for any specific process or welding current. Such data can be found in publications referenced below or by a web search. The ranges of approximate deposition rates for the commoner arc welding processes are listed in Table 1.

Table 1 Indicative deposition rates - arc welding processes

Welding ProcessDeposition Rate kgs/hr

min max

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MMA 0.4 5.5MAG 0.6 12FCAW 1.0 15Single wire SAW 3 16To obtain an accurate figure for the specific parameters to be used is a relatively simple exercise. Weighing a plate, depositing weld metal using the required parameters on this plate for a fixed time and then re-weighing the plate will give an accurate figure that may be used for estimating purposes.

There is one golden rule for minimising the cost of making a weld and, whilst this may seem to be self-evident, it is worth repeating: deposit the minimum amount of the highest quality weld metal with the largest gauge electrode or wire at the highest current in the shortest possible time. This is obviously the ideal and can seldom be achieved in practice because of limitations on heat input, access etc.

The implications of applying the golden rule are:

1. To deposit the minimum amount of weld metal the designer, aided by the welding engineer, must select the smallest weld preparation that is capable of providing the required weld quality. If the included angle is too narrow then lack of side wall fusion is a possibility with the consequent costs of repair; too wide an angle is wasteful in terms of deposited weld metal. Remember, though, that the cost of providing a weld preparation (by flame-cutting, edge planning, milling etc) must also be included in any costing exercise as must the cost of assembly. Machined weld preparations are more accurate than flame cut edges and this may result in faster set-up times and a reduced weld repair rate.

It may be possible to use a square edge preparation by using the deep penetration characteristics of some of the welding processes; electron beam and laser welding are the best examples of this technique. Plasma-TIG and activated flux TIG can penetrate up to 10mm in a single pass; the 'finger' penetration of spray transfer MAG welding can penetrate up to 6mm and a submerged arc weld can penetrate up to 15mm. There is also the benefit when using a square edge preparation in that the consumption of filler metal is substantially reduced, the bulk of the weld metal being provided by the parent material.The final option on reducing costs when butt welding is for the designer to specify a partial penetration joint. The most expensive weld pass in any full penetration butt weld is the root pass and if this can be eliminated by using partial penetration joints then substantial savings can be made. However, the decision to use partial penetration welds should not be taken lightly but only if service conditions permit the presence of a large crevice at the weld root. The designer will therefore need to consider whether fatigue, creep, corrosion etc are likely to occur and must clearly specify where the joints are permitted and the minimum acceptable weld throat.

2. Depositing the highest quality weld metal infers that the weld repair rate will be reduced. Repair weld metal is very costly, particularly if the unacceptable defects are detected late in the fabrication programme; perhaps after final assembly where access is difficult or after post weld heat treatment. Accurate weld preparations and fit-up, easy access for the welder, welds made in the flat position and well trained welders will all help to minimise the weld repair rate.

3. Depositing weld metal with the largest electrode or wire at the highest current will obviously give the highest weld deposition rate and shortest joint completion time. The deposition rate figures in Table 1 give the minimum and maximum deposition rates at minimum and maximum welding currents. As an example, a 1.2mm diameter MAG wire at 120amps will deposit around 1.2 kgs/hr, at 380amps around 8 kgs/hr. To enable high welding currents to be used the item must be placed in the flat position and there must be easy access for the welder. One benefit of using the high welding currents is that the number of weld runs to fill the joint will be reduced and this, in most circumstances, will result in less distortion than a large number of low current weld passes. Remedial work to correct distortion can

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therefore be reduced. A further benefit when welding the ferritic steels is that high current and therefore high heat input may allow any preheat to be reduced or eliminated entirely.

However, there are limitations to this approach to improving productivity. If achieving high toughness is a factor then it is likely that heat input will need to be controlled when welding the ferritic steels, placing a limit on the welding current and travel speed. High welding currents also imply a large, fluid weld pool and it may not be possible to control this pool when welding in any other than the flat position - for example, MAG welding cannot be performed using spray transfer (high welding current) in the vertical position due to the absence of a flux to hold the pool in place. Using a manual process at such high currents also results in increased welder fatigue resulting in a reduced duty cycle. A solution to this problem is to mechanise or automate the process.

To achieve the most cost effective solution to producing a welded structure is therefore not simply to increase duty cycle or deposition rate but to consider all aspects of fabrication from the design stage to final inspection, involving all members of the team from designer to welder.