paaf

333

Weld repairs

naldine peS eqbsehe T

ationh the oany petime aon suffe airc

Basically, the aeronautical projects should take into account thedifculties of transporting a load against the gravity force duringtake-off and ight, and discharge it in an efcient way, with mini-mum cost and maximum safety, because failures in any of thesestages will implicate catastrophic accidents, involving human lives[3]. In general, structural failures during ight are attributed to

accomplished by adoption of prevention methods with high-qual-ity periodic inspections associated to damage tolerance philoso-phy, for taking into account its effect on the fatigue life of astructural component [4,5,1012]. Such issues will request exten-sive investment, planning, researches and development on repairmethods and maintenance procedures towards assure the safeand continued airworthiness.

For several aircraft models (e.g. agricultural, military trainingand acrobatic) the most solicited and repaired component is onethat supports the motor, called Motor-Cradle (Fig. 1) [13]. This

* Corresponding author. Tel.: +55 12 3123 2865; fax: +55 12 3123 2852.E-mail addresses: pereira@feg.unesp.br, marcelino.nascimento@gmail.com (M.P.

International Journal of Fatigue 32 (2010) 12001209

Contents lists availab

u

l sNascimento), voorwald@feg.unesp.br (H.J.C. Voorwald).with the aircraft or anything attached thereto, or in which the air-craft received substantial damage (damage or structural failurethat adversely affects the structural strength, performance, oright characteristics of the aircraft, and which would normally re-quire major repair or replacement of the affected component).Based in such denition, the accident rates (i.e., accidents per mil-lion departures) have been at 1.2, or 12,000 accidents, in the occi-dental world [2]. In the search for a zero accident rates, the ight-safety has been the main concern of the aeronautical authorities allover the world.

Since the aircrafts becomemore complex, the environmental ef-fects assumed great importance. As a result of older aircrafts yingnowadays, problems such as stress-corrosion cracking, corrosion-fatigue (simultaneously or separately) and wear are also expectedto occur [4,5,9,10]. Due to several aggressive environments inwhich the aircrafts are subjected, particularly marine, corrosionis the most important factor of maintenance and inspection forthe aeronautic sector [4,5,9,10]. Hence, corrosion is undoubtedlya real and critic issue acting on aircrafts, owing with the timeeven on those not operational. Usually, the corrosion control isFlight-safety

1. Introduction

According to ight-safety foundmeans an occurrence associated witwhich takes place between the timewith the intention of ight until suchdisembarked, and in which any persjury as a result of being in or upon th0142-1123/$ - see front matter 2010 Elsevier Ltd. Adoi:10.1016/j.ijfatigue.2009.12.017[1], Aircraft accidentsperation of an aircraft,rson boards the aircrafts all such persons haveers death or serious in-raft or by direct contact

aerodynamic overloads or fatigue of materials, as a consequenceof inadequate project or any notch produced during manufacturingor maintenance of aircrafts [47]. Since the catastrophic accidentsinvolving the English model Comet in the 1950s, the fatigue pro-cess has been the most important project and operational consid-eration for both civil and military aircrafts [4,5,8]. On the otherhand, many fractures of materials are also caused by corrosion asa consequence of aggressive environment.Fatigue behaviorCorrosionConsiderations on corrosion and weld reof a steel structure critical to the ight-s

Marcelino P. Nascimento *, Herman J.C. VoorwaldDepartment of Materials and Technology State University of SoPaulo, UNESP/FEG/DMT

a r t i c l e i n f o

Article history:Received 30 July 2009Received in revised form 24 December 2009Accepted 31 December 2009Available online 7 January 2010

Keywords:AISI 4130 steel

a b s t r a c t

The aim of this study is to arepairs on the reverse benight-safety. The tests werby means of a SCHENK PWroom temperature. It was odue to the corrosion and t

International Jo

journal homepage: www.ell rights reserved.ir effects on the fatigue strengthety

, Ariberto Pereira da Cunha Ave., 12516-410, Guaratinguet City, So Paulo State, Brazil

yze the effects of corrosion and successive tungsten inert gas (TIG) weldingg fatigue strength of AISI 4130 steel used in components critical to therformed on hot-rolled steel plate specimens, 1.10 mm and 1.60 mm thick,uipment, with load ratio R = 1, constant amplitude, 30 Hz frequency andrved that the reverse bending fatigue strength of AISI 4130 steel decreasesIG welding and re-welding processes.

2010 Elsevier Ltd. All rights reserved.

le at ScienceDirect

rnal of Fatigue

evier .com/locate / i j fa t igue

ionacomponent presents a geometrically complex structure made fromAISI 4130 tubular steel of different dimensions and TIG welded inseveral angles [13,14]. For the Brazilian aircrafts T-25 Universaland T-27 Tucano, for example, besides supporting the motor in bal-ance, the motor-cradle also maintains the nose landing gear xedat the other extremity. Since the motor-cradle is a component crit-ical to the ight-safety, the aeronautic standards are extremely rig-orous in its manufacturing, by imposing a zero-defect index onthe nal weld seam quality (Safe-Life philosophy), which is 100%inspected by non-destructive testing/NDT [1315]. For this reason,welded aeronautic structures are frequently subjected to succes-sive repairs in accomplishment to current standards. As a conse-quence, components approved by NDT may contain a historicrecord of welding repairs whose effects on their structural integrityare not computed. In addition, these structures are also submittedto weld repairs along their operational life, turning this questionmore complex.

As a part of this research-work, an investigation on 157 motor-cradles fracture reports indicates that all of them occurred atwelded joints as a result of fatigue cracks, reducing the Time-Be-fore-Fail from 4.000 h to 50 h of ight [13]. Motivated by highfracture incidence of this particular component, an extensive re-search program to evaluate the manufacturing and maintenanceweld repair effects on the structural integrity, mechanical proper-ties and microstructural changes has been conducted [13].

Although the maintenance repairs of parts and components area multibillion-dollar industry [16] and that, particularly for the

Fig. 1. Motor-cradle assembly in a T-25 Universal Brazilian aircraft.

M.P. Nascimento, H.J.C. Voorwald / Internattransport sector, the welding repairs are an essential and fre-quently used process [17], few papers approaching this issue havebeen published, being all of them on either aged and degeneratedmaterials [18,19] in petrochemical, offshore and power industries[20], by means of nite element method FEM - [2123].

In this paper, special emphasis was attributed to a standardizedweld repair procedure, widely employed during manufacture ofwelded aeronautic structures and characterized by removing thepreviously defective weld bead and by applying a new weld seamby means of the gas tungsten arc welding/GTAW (or tungsten inertgas/TIG) process with ller metal.

The aim of this study is to analyze the effect of the corrosiveprocess and the successive TIG welding repairs on the reversebending fatigue strength of AISI 4130 steel specimens made fromhot-rolled steel plates, 1.60 mm and 1.10 mm thick, respectively.Analyses of microstructure, microhardness and residual stressesfrom the base-material (BM), heat-affected zone (HAZ) and weldmetal (WM), as well as the effects of the surface roughness andweld bead geometry complemented this study.

In the sequence, a brief comment on the current weld repairprocedures is presented.1.1. Current Welding Repair Procedure and Applications

When defects (such as porosity, impurity, undercut) arefound in the welding joints, the main international standards(e.g. API, AWS, BSI and ASME) request that the defective weldedjoints are corrected. However, in a general form, none of theabove standards make mention to number of permissiblerepairs.

In general, repair welding usually involves the completeshut-down of the equipment, the removal of all dead loadsand the carrying out welding to original manufacturing stan-dards [24]. Two typical procedures of weld repairs, complete re-pair and partial repair, include mechanics removal of thedamaged material and the restoration of the geometry andintegrity of the component, followed for a post-welding heattreatment (PWHT), which will depend mainly on type andthickness of material [21]. PWHT has as a function to reduceor eliminate residual stresses and to temper the metallurgicalstructure of heat-affected zone (HAZ). However, PWHT is veryexpensive, time consuming and, depending on local, difcultor impracticable.

An important application of weld repair techniques is to avoidPWHT, generating HAZ with structures of ne grains [24]. Suchan objective can be reached by minimizing the size of original grainin HAZ, by minimizing the thermal contribution during the weld-ing process or by rening the size of the initial coarse-grain regionof HAZ [24]. For this purpose, two welding repair techniques areemployed, viz:

i. Half-bead welding (HBW): the beads are removed beforewelding the next layer.

ii. Temper-bead welding (TBW): the subsequent weld beadpartially does heat-treat the previous passes, after that thedeposited weld layer is removed.

Both techniques involve removal of the upper passes and theyare employed to avoid PWHT, to reduce or eliminate residual stres-ses, to avoid hydrogen cracking, to improve fracture toughness andto reduce welding cost and time [24].

For the aeronautic sector, in accordance with IIS/IIW 956-87Doc [25], when repairing a weld seam in critical aeronautic compo-nents is requested that the defect is also located and removed. Theremoval process should be always carried out from the side thatpropitiates the smallest weld material loss. Soon after, to maintainthe uniformity of the deposited metal along the weld seam, the re-pair is applied [25].

However, the weld bead removal can be uneconomical,impracticable or even impossible in certain engineering applica-tions due to urgency or emergency, difcult access to the com-ponent or appropriate equipments. As a consequence, along theuseful life of the aircraft many non-standardized weld repairsare carried out by putting upon several weld bead, without re-moval of the previous pass. The subject to be answered up iswhether or not the weld repairs are a viable and effective pro-cedure, capable of extending the in-service life of componentsand structures. At present, due to the increase of residual stres-ses, the distortion in the geometry and deterioration of themicrostructure caused by the thermal cycles, the effectivenessof the weld repairs on cracked structures is not well under-stood and investigated. Differently to the considerable data vol-ume on the effects of original weld on fatigue life for steelsand aluminum from the literature, data on welds repairs arescarce. In this context, the availability of experimental data

l Journal of Fatigue 32 (2010) 12001209 1201on weld repair effects on the fatigue behavior may be very use-ful in determining inspection intervals in high-responsibilitywelded structures.

2. Material and methods

2.1. Material

For the present research-work at welded specimens from hot-rolled AISI 4130 aeronautic steel, 1.10 mm and 1.60 mm thick,were used. The chemical compositions (wt%), from the base-mate-rial and from the weld (ller) metal, are presented in Table 1 (Fe inbalance). The mechanical properties obtained from smooth atsamples and original welded specimens (OR) with central weldseam crossed to the hot-rolled plate direction , are indicated inTable 2. The hot-rolled plates presented 65 HRA in the as-re-ceived condition.

The monotonic tensile tests were performed in accordance withASTM E 8 M by means of a servo-hydraulic INSTRON test machine,applying 0.5 mm/min displacement rate and a preload equal to0.1 kN.

only one weld single-pass was required. For re-welding process, amanual grinding machine capable to reach 22,000 rpm was used toremove the previous weld metal. The heat-input applied was keptconstant for all the welded and re-welded specimens.

2.3. Corrosion process

To analyze the effects of the corrosive process on the fatiguebehavior of AISI 4130 steel, specimens were submitted to atmo-spheric corrosion (urban environment), accelerated by means ofconstant wetting during thirty days. Periodically, the specimenswere moved so that the all the surfaces were uniformly corroded.

2.4. Reverse bending fatigue tests

For experimental bending fatigue tests, welded and non-weldedat specimens were manufactured according to SCHENCK modelPWS requirements (Fig. 2), following the LT direction of the hot-rolled plate. The specimens were fatigue tested upon a sinusoidalload, constant amplitude, load ratio R = 1, at 30 Hz frequencyand room temperature. The supercial average roughness, ob-tained by means of a Mitutoyo 301 equipment, with cut-off equalto 0.8 mm 5 mm, was Ra = 0.24 lm 0.16 lm.

2.5. Microstructural and microhardness analyses

For microstructural analysis, a Nital 2% chemical etching wasapplied during 5 s. Vickers microhardness measurements were ob-tained at 0.0254 mm intervals throughout the regions under anal-ysis (base-material, HAZ regions and weld metal), applying a 1 Nload.

S

00000

d st

ra s

1202 M.P. Nascimento, H.J.C. Voorwald / International Journal of Fatigue 32 (2010) 12001209Table 1Chemical compositions (wt%).

Composition (wt%) C Mn Pmax

Specied BMa 0.280.33 0.400.60 0.035Specied WMb 0.280.33 0.400.60 0.008Weld (ller) metal 0.30 0.50 0.004Plate #1.10 thick 0.33 0.53 0.010Plate #1.60 thick 0.28 0.49 0.009

a AMS 6457 B turballoy 4130 steel.b AMS T 6736 A (2003) for chromiummolybdenum (4130) seamless and welde

Table 2Mechanical properties obtained from the welded and non-welded specimens.

Specimens Yielding (0.2% offset) Ultimate strength Ruptu2.2. Welding and re-welding procedures

The most commonly employed welding process for manufac-turing of aeronautical structures is tungsten inert gas (TIG), orgas tungsten arc welding (GTAW), which is appropriate to weldthin thickness materials and to allow the necessary variable con-trol, resulting in high-quality and defect-free weld beads. The TIGwelding process was carried out in accordance with the BrazilianAeronautic Industry EMBRAER NE 40056 TYPE 1 Standard(for components critical to the ight-safety), with a protective99.95% purity argon-gas and ller metal AMS 6457 B Turballoy4130. A Square Wave TIG 355 Lincoln equipment was manuallyemployed by an expert aeronautic welder. All the welding param-eters were controlled, and the main ones are indicated in Table 3.Also, all the welded joints were subjected to X-ray non-destructiveanalysis by the Brazilian Aerospace Technical Centre (CTA/IFI),which proved the acceptable quality of the welds, according toMIL-STD-453, EMBRAER NE-57002 and ASTM E-390 standards.

The welding direction was always perpendicular to hot-rollingprocess (direction) of the plate. Before welding, samples werecleaned with chlorinated solvent to oxide removal and xed on abacking bar, to avoid contamination and porosity in the weld root.After the welding/re-welding process neither subsequent heattreatment to residual stresses relief nor subsequent removal ofthe weld bead was carried out, in order to simulate the real condi-tion of the original aeronautic structure. Due to the plate thickness,Base-metal 746 21 843 9 655 26Welded 671 20 778 17 643 26max Si Mo Cr Cu

.040 0.150.30 0.150.25 0.801.10 0.10

.008 0.150.35 0.150.25 0.801.10 0.10

.003 0.25 0.18 0.91 0.042

.003 0.28 0.17 1.04 0.02

.004 0.24 0.18 0.93 0.02

eel tubing of aircraft quality.

tress Elongation (25 mm length) Yielding/ultimate strength ratio

Table 3TIG welding parameters.

Direct current DCEN

Welding position FlatWelding voltage 12 VWelding current 30 AWelding average speed 19.0 cmmin1

Pre heating NONEFlow rate 412 L min1

Theoretical heat input 1.2 kJ cm1

Filler metal diameter 1.6 mm9.80 1.87 0.88 0.023.81 0.26 0.86 0.01

mens were removed by electrolytic polishing with a non-acid

integrity, by providing high capacity of plastic deformation andconsequent margin of safety against fracture [27]. Yet, all themonotonic specimens tested, including those grinding machined,fractured at the sub-critical HAZ (SCHAZ)/base-material interface(strength-overmatch), despite any stress concentration originatedat the weld toe, as illustrated in Fig. 3. This means that the inu-ence of the microstructural variations along HAZ is more deleteri-ous for tensile strength than the geometric features induced by thewelding process. That is, for the monotonic test the material wasnot sensitive to stress concentration at the weld toe.

3.2. Reverse bending fatigue tests

3.2.1. Effect of the corrosion processFig. 4 presents the SN (stress vs. number of cycles) curves for

corroded and non-corroded AISI 4130 steel specimens. For seekof comparison with the welded specimen, it was added the fatiguecurve obtained with base-material specimens (BM) 1.10 mm thick.Considering that for all the specimens just the thickness was al-tered (1.601.10 mm), it is possible to verify that the higher the

80

40 22

50

20

8,5 (4x)

M.P. Nascimento, H.J.C. Voorwald / International Journal of Fatigue 32 (2010) 12001209 1203solution.

3. Results and discussion

3.1. Monotonic tests

First of all, it is important to pay attention to the high mechan-ical strength values and reasonable ductility from the hot-rolledAISI 4130 steel plate (as presented in index 2.1). It is also interest-ing to observe the decrease of all that mechanical properties afterthe rst TIG welding application on AISI 4130 steel specimens, par-ticularly the elongation (3.8% average), typical of fragile material.However, for both welded and non-welded conditions the yield2.6. Residual stresses determination

The residual stresses eld induced by welding and re-weldingprocesses was determined by X-ray diffraction method, using theRaystress equipment (whose features are described in [26]) withcouple exhibition, u goniometer geometer, two anodes of chromeCrka radiation and registration of {2 2 1} diffraction lines, tensionsource of 25 kV, current of 1.5 mA, X-ray convergence angle of 50.The accuracy of the stress measurements was Dr = 20 MPa. In or-der to obtain the stress distribution by depth, the layers of speci-

110

Fig. 2. Flat bending fatigue specimen (mm).tensile stress/maximum tensile stress ratio (ry/rm) was around0.8, which is an appropriate value for structures as the aeronauticalones [27]. Historically, values from 0.80 to 0.86 have been consid-ered appropriate for specication, project and analysis of structural

Fig. 3. Specimens fractured by monotonic tension tests.thickness the lower the fatigue strength. It is in accordance withMurakami [28,29], who mentions that the larger specimens thehigher amount of existent defects. Makkonen [30] also mentionsthat there is a great amount of microcracks previously nucleatedin specimens submitted to cyclic loads, whose population in-creases when their dimensions increase.

From Fig. 4 it is also possible to verify that the fatigue curves di-verge in the short fatigue life range/SFL (

crack tip accelerates the corrosive process and vice versa), the mainfactor controlling the fatigue behavior presented is related to sur-face roughness of the specimens tested. In fact, the average surfaceroughness measured provided Ra = 0.24 lm 0.16 lm for non-cor-roded specimens and Ra = 1.95 lm 0.18 lm for those corroded.The corrosive process generated a Grant No. of pits, which gave riseto local stress concentration and reduced the number of cycles forcrack nucleation. Additionally, the local stress concentration canalso reduce the stress intensity factor threshold (DKth) and allowsthat the small cracks grow [1012]. Therefore, it is necessary toprotect the surface of aeronautic structures against the aggressiveenvironment acting on an aircraft.

3.2.2. Effect of the TIG welding/re-welding processFig. 5 presents the SN curves of welded and re-welded speci-

mens. The horizontal line indicates the nominal stress value(rn 247 MPa), which corresponds to yielding stress divided bythe safety-factor equal to 3 (for welded components critical tothe ight-safety), in accordance with the EMBRAER NE 40056

metric alteration was hoped to be low (as can be seen in Table 5ahead). Thus, the increase of the notch sensibility (fatigue-notchfactor) with the second weld repair (2R) was also due to the micro-structural variations resulting of successive heating/cooling cyclesimposed to the material. Therefore, based on the results obtainedand considering that the welding repair is necessary, it can beadmitted only one weld repair (1R) during fabrication of the criti-cal to the ight-safety structures, without compromising their re-verse bending fatigue strength. On the other hand, being

1204 M.P. Nascimento, H.J.C. Voorwald / InternationaTYPE 1 Standard.From Fig. 5, one can observe the signicant decreasing in bend-

ing fatigue strength of the AISI 4130 steel, whose endurance limitwas about 37% of its yielding stress (ry). However, the endurancelimit is still above the nominal stress, rn (horizontal line). Never-theless, a subsequent reduction in bending fatigue strength forwelded specimens is observed in comparison with the base-mate-rial specimens (BM), mainly in the short fatigue life (SFL) range. Inthe long fatigue life (LFL) range, it is possible to verify that theendurance limit is located close to the horizontal line, correspond-ing to specied nominal stress (rn). On the other hand, with therst weld repair (1R), no-subsequent reduction in bending fatiguestrength is observed. For both original (OR) and rst time re-welded (1R) specimens, one can observe practically the same fati-gue behavior. However, the second weld repair (2R) resulted in adecreasing of the fatigue strength, with the endurance limit belowthe nominal stress value, rn (horizontal line). Since the aircrafts aresubmitted to high fatigue cycles during ight, as a result of abruptmaneuvers, wind bursts, motor vibration and helixes efforts, it isnot recommended or favorable to the ight-safety the second weldrepair (2R) application. This behavior is also due to the increase involume of the weld bead with consequent increase of the stressconcentration factor at the weld toe. In addition, the re-weldingprocess can increase the HAZ dimensions and its coarse-grain re-

105 106 107

150

200

250

300

350

400

450

500

SAFETY FACTOR=3

37% y

32% y

20% y

MAX

IMU

M S

TRES

S [M

Pa]

CYCLES [N]

BASE-MATERIAL (#1.10mm)ORIGINAL WELDING1 REPAIR OF WELDING2 REPAIRS OF WELDINGCORRODED BM (#1.60mm)Fig. 5. Reverse bending fatigue SN curves for base-material and (re-)weldedspecimens (1.10 mm thick).gion (CGHAZ), which is located exactly at the weld toe and charac-terized by low fracture toughness and considerable hardness (ascan be veried in Table 4).

From Fig. 5, it is also possible to verify the larger scattering ofthe fatigue results obtained with both welded and re-welded spec-imens. This is due to volume variations of the deposited weld me-tal, the higher heating/cooling rate and the uneven stressconcentration values at the weld toe notch. Fig. 5 also shows thatthe fatigue strength reduction due to welding process tends to bemore pronounced when the number of cycles is also reduced(divergence between BM and OR fatigue curves in SFL). This obser-vation is very important taking into account the efforts due tolanding operations (typical for SFL regime) along the operationallife of the aircraft. On the other hand, it is observed that the diver-gence between both endurance limits (107 cycles) was low, i.e.:from 37% ry for base-material to 32% ry after welding (close torn). In the same way, this is a very important result, taking into ac-count the abrupt maneuvers or wind bursts during ight (typicalfor LFL regime). This implicates that special cares should beadopted during design and, particularly, maintenance of weldedaeronautical structures against any accidental tool marks capableto introduce local stress concentration on the structure. Since forbending loads the maximum stress occurs at the surface, the fati-gue resistance is sensitive to any geometric change of material(notch-sensibility).

From Fig. 5, comparing the rst (1R) and second weld repairs(2R) fatigue curves in the SFL (

bined, are considered the main geometric factors controlling thefatigue behavior of welded components and structures [13,31]. Inaddition, the great scattering on the geometric factors measuredis in consonance with the great scattering of the fatigue resultspresented by the welded specimen groups. Yet, from Table 5, wecan observe: the increase of the stress concentration factor (Kt) atthe weld toe notch with the successive welding repairs; the largestextension of HAZ due to the successive heat-input applied and,consequently, the increase of the CGHAZ as well; the effect of bothweld reinforcement (T) and weld root dimensions, again, on the an-gle (a) and radius reduction and, consequently, on the bending fa-tigue strength of welded specimens. On this subject, it is important

Angle (a) Radius (mm) HAZ (mm) Kt (Eq. (1))

141.9 14.4 1.03 0.36 2.89 0.25 1.290146.1 8.6 0.75 0.21 3.11 0.12 1.305138.6 7.5 0.93 0.40 3.22 0.23 2.178

1 32e-notch factor.

Fig. 7. Crack initiation at the weld toe of a non-fractured fatigue specimen (500).

ional Journal of Fatigue 32 (2010) 12001209 1205addition, it becomes necessary lower inspection intervals for awsas-crack determination at the weld joints of structuralcomponents.

Finally, although it is not appropriate to directly compare bothfatigue behaviors of welded specimens with those corroded due totheir different thickness, it is interesting to observe the similar del-eterious effect caused by the corrosive process on the fatiguestrength of AISI 4130 steel.

Table 4 presents the microhardness values measured in thethree interest areas. Firstly, one can observe the high microhard-ness values for weld metal and CGHAZ as a consequence of TIGwelding and re-welding process. This is due to high cooling speedassociate to the low heat-input applied, which resulted in martens-itic structures. Since the heat input was kept constant for all the(re-)welding process, it is important to mention that just the weldreinforcement was removed. Consequently, the different microh-ardness values veried were due to remaining weld material vol-umes (e.g. weld root), which unevenly affected the cooling speedof the weld metal. Thus, from Table 4, one can observe that for both

Table 5Geometry factor values from the weld bead (in accordance with Fig. 6).

Group W (mm) T (mm) Root (mm)

OR 3.75 0.35 0.89 0.20 0.77 0.191R 4.49 0.33 1.17 0.32 0.82 0.272R 4.80 0.23 0.96 0.20 0.79 0.30

Kf 1 Kt11a=r ;where: a = 0.1659 (Petersons material parameter for steel); r = notch-root; Kf = fatigu

RADIUS

ROOT

HAZW

T

Fig. 6. Geometry factors of the welded joints (values presented in Table 5).

M.P. Nascimento, H.J.C. Voorwald / Internatthe original weld (OR) and the rst welding repair (1R) the microh-ardness values were close and coherent with each other. This canexplains the bending fatigue behavior presented in Fig. 5. However,it is also important to pay attention to the great dispersion ofmicrohardness results in both CGHAZ and weld metal (standard-deviation). It is also possible to observe the highest microhardnessvalue from the original weld metal than for all the other condi-tions. This implicate that probably the second welding repair pro-moted the tempering of the previous microstructure, but with noimprovement on their bending fatigue strength, as veried inFig. 5. After the second welding repair it is also possible to observehigher microhardness value for the CGHAZ, implicating a grain sizereduction in that region, but again with no improvement on thebending fatigue strength, as illustrate in Fig. 5. It is well known thatthe lower the grain size the higher the material toughness. In thesame way, the higher the grain size, the lower the hardness/mechanical strength.

Considering that the weld prole affects the fatigue resistanceof a welded structure, Table 5 and Fig. 6 present the main geomet-ric factors that compose a weld joint and their corresponding val-ues, obtained by image analysis tools. Thus, from Table 5 and Fig. 6,it is possible to conrm all the results presented in Fig. 5 on the ef-fect of welding repairs on the bending fatigue behavior of AISI 4130steel. Therefore, it can be observed that the welding repairs re-duced both angle (a) and radius at the weld toe notch that, com-Nital 2%.Fig. 8. Macrography of a fatigue specimen fractured by reverse bending loads. Nital2%.

Fig. 9. Secondary crack propagation from the fracture surface of a reverse bendingfatigue specimen (500). Nital 2%.

to mention that all the welded specimens fractured at the weld toenotch, as illustrated in Figs. 7 and 8. In fact, Fig. 7 illustrates a crackin a non-fractured specimen, which initiated at the weld toe notchand propagated through grain boundary into weld metal (inter-granular crack).

Fig. 8 presents the macrography of a fractured specimen thatwas submitted to two welding repairs (2R), in which we can ob-serve: (a) the weld metal region with well delineated equiaxialcoarse-grains; the HAZ extension with the sub-regions, (b)coarse-grain/CGHAZ, (c) ne-grain/FGHAZ, (d) inter-critical/ICHAZand (e) sub-critical/SCHAZ near to A1 line (lower critical tempera-

ture); as well as the welded joint irregularity as a consequence ofthe deformation generated by the reverse bending loads applied.

Fig. 9 presents the transgranular secondary crack, originatefrom the fracture surface of the specimen in Fig. 8. This secondarycrack implicates that the reverse bending fatigue specimens mayhave been subjected to either Modes I and II or Modes I and III ofloading combination (or all of them), turning the laboratorial testmore severe due to friction process by both free crack surfaces,accelerating its propagation. In fact, bending loads also introduceshear stresses in the specimens (Mode II of fracture) and/or tearingstress (Mode III of fracture) whether crack initiation occurs at theedge of the specimen, as illustrated in Fig. 10. On this subject, Veidt[33] mentions that the radius in the notch root has little effect onthe stress intensity factor in Mode II of fracture.

Figs. 11 and 12 present the base-material, HAZ and weld metalmicrostructures for all the proposed conditions. Fig. 11a shows thenormal products of transformation from austenite, i.e. ferrite andperlite. From Fig. 11bd the beginning of the transformation fromperlite to austenite is observed (to martensite, upon subsequentcooling) along the SCHAZ/ICHAZ regions (A1 line).

Fig. 12 presents the microconstituent of the weld metal (origi-nal, one and two weld repairs) and respectives CGHAZ, basicallyconstituted by martensite, which is a very hard and fragile micro-structure. As aforementioned, the higher the hardness the higherthe resistance to fatigue crack initiation.

Fig. 13 presents the residual stresses eld in the weld metal,HAZ and base-material of all the conditions analyzed. It is wellknown that residual stresses are present in welded components[22,3436] and have great effect on fatigue crack nucleation andpropagation [34]. Thus, from Fig. 13, one can observe that the

Fig. 10. Edge fatigue crack origin and propagation for a fractured specimen of base-material, 1.60 mm thick.

1206 M.P. Nascimento, H.J.C. Voorwald / International Journal of Fatigue 32 (2010) 12001209Fig. 11. Base-material and base-material/HAZ transition microstructures: (a) Base-matwelding repair; and (d) ICHAZ of the second welding repair. Nital 2%. (F = Ferrite; P = Peerial (typical); (b) SCHAZ-ICHAZ of the original weld; (c) SCHAZ-ICHAZ of the rstrlite; M = Martensite).

TIG (re-)welding process induced high compressive residual stres-ses for all conditions upon analysis, as weld metal as in HAZ andbase-material (up to 20 mm from the fusion line). Such residualstresses were relieved internally in some point along the lengthof the specimens, because no deformation was veried on the sam-ples. Yet, it is observed that all the residual stress proles pre-sented similar tendency, i.e., maximum values in weld metal(600 MPa for original weld; 450 MPa for one weld repair and330 MPa for two weld repairs), followed by HAZ (400 MPa fororiginal weld; 75 MPa for one weld repair and 50 MPa for twoweld repairs) and last base-material up to 20 mm far-away fromthe fusion line (300 MPa for original welding; 160 MPa forone weld repair and 100 MPa for two weld repairs).

Many factors might have contributed to compressive residualstresses induced by the TIG (re-) welding process, viz: (i) austen-itemartensite transformation (which generate up to 4% increasein volume of material) and which initiate at the surface due tohigher cooling speed; (ii) the thin thickness and good mechanicalproperties in high temperature of the base-material, particularlylow deformation in high temperature; (iii) subsequent contractionin the core of the weld metal (after the martensitic transformationat the surface) due to relatively lower cooling speed in that region;(iv) constraint to natural expansion of the weld metal and HAZ bythe base-material volume around; and (v) the clamps applied atthe extremities of the specimens during the (re-)welding process,etc. However, some aspects get the attention from the Fig. 13 as:

M.P. Nascimento, H.J.C. Voorwald / International Journal of Fatigue 32 (2010) 12001209 1207Fig. 12. Microscopic analysis of the weld metal and CGHAZ: (a) original CGHAZ; (b) oriwelding repair; (e) CGHAZ of the second welding repair; and (f) weld metal of the secoginal weld metal; (c) CGHAZ of the rst welding repair; (d) weld metal of the rstnd welding repair. Nital 2%.

loading in aluminium alloys of aeronautical applications. Ph.D. Thesis, COPPE/

[7] Latorella KA, Prabhu PV. A review of human error in aviation maintenance and

ionathe high values of compressive residual stresses, which were high-er in the weld metal than HAZ; all the compressive stress eldswere still compressive up to 20 mm distance from weld fusion line.

It is well known that the residual stresses largely affect the fa-tigue behavior of components, and that compressive residual stres-ses are benet towards inhibit the crack nucleation. Thus, theresults presented in Fig. 13 are in accordance with the fatiguebehavior of specimens with original weld (OR), as well as thosewith one (1R) and two (2R) weld repairs presented in Fig. 5. Inaddition: the higher the hardness, the higher the fatigue strengthas well. Consequently, from Table 4 and Fig. 5, it could be observedbetter fatigue behavior for the originally welded (OR) specimens incomparison with the one (1R) weld repair condition. However,considering the fatigue crack propagation stage, it is well knownthat the increase of the hardness increases the propagation rate(da/dN). In this context, the existence of compressive residualstresses highlight the stress concentration effect induced by theweld geometry on the fatigue strength reduction veried. That is,although a mechanical component possess high compressive resid-ual stresses, their effects or benets will be minimized (or annulled

WELD METAL FGHAZ BM (20mm WM)-650

-600

-550

-500

-450

-400

-350

-300

-250

-200

-150

-100

-50

RES

IDU

AL S

TRES

SES

[MPa

]

REGIONS OF THE WELDED JOINTS

1 WELD REPAIRORIGINAL WELD2 WELD REPAIRS

Fig. 13. Residual stress proles of all the conditions tested.

1208 M.P. Nascimento, H.J.C. Voorwald / Internatin case the stress relief) when this component presents a stressconcentrator (geometric factor) that reduces the number of cyclesnecessary to fatigue crack initiation. Thus, it is possible that thegeometric stress concentration factor, which is located at the weldtoe/fusion line/CGHAZ region, overcomes the compressive residualstress eld induced by the (re-)welding process, relief it. Certainly,the compressive residual stresses eld as far as possible delayedthe fatigue crack nucleation and propagation by reducing the stressintensity factor (DK), as mentioned by Wei & Chen [34].

4. Conclusions

Motivated by high fracture incidence at welded joints of a spe-cic and critical to the ight-safety component, called motor-cra-dle, experimental reverse bending fatigue tests on welded and re-welded specimens were carried out. Based on the results obtained,the following conclusions may be drawn:

1. The AISI 4130 steel possess good mechanical properties, but lowelongation in the as-received condition (not heat-treated).

2. All the monotonic specimens tested, including those grindingmachined, fractured at the sub-critical HAZ (SCHAZ)/base-material interface (strength-overmatch), despite any stress con-centration originated at the weld toe.inspection. Int J Ind Ergonom 2000;26:13361.[8] Bhaumik SK, Sujata M, Venkataswamy MA. Fatigue failure of aircraft

components. Eng Fail Anal 2008;15:67594.[9] Carpenter M. Managing the eet: materials degradation and its effect on aging

aircraft. Amptiac, News Lett 2001;4(5):719.[10] DOT/FAA/AR-00/22. US Department of Transportation, Federal Aviation

Administration, Corrosion and Corrosion Fatigue of Airframe Materials, nalreport, July 2002. 79p.

[11] Wang QY, Kawagoishi N, Chen Q. Effect of pitting corrosion on very high cyclefatigue behavior. Scripta Mater 2003;49:7116.

[12] Murtaza G, Akid R. Corrosion fatigue short crack growth behaviour in a highstrength steel. Int J Fatigue 1996;18:55766.

[13] Nascimento MP. Effects of TIG welding repair on the structural integrity of AISI4130 aeronautical steel. In: Ph.D. Thesis in mechanical engineering. StateUniversity of So Paulo/UNESP-FEG, Brazil (in Portuguese); Code CDU 620.92,2004. 240p.

[14] Nascimento MP, Voorwald HJC. An evaluation on the fatigue crack growth inre-welded AISI 4130 aeronautical steel. In: Eighth international fatiguecongress/FATIGUE 2002, vol. 5, StockholmSweden; 2002. p. 346372.

[15] Nascimento MP, Ribeiro RB, Voorwald HJC. Fatigue crack growth in re-weldedAISI 4130 high strength steel. In: Fourteenth European congress on fatigue/ECF14, vol. 2, Cracow, Poland; 2002. p. 13116.

[16] Vern Sutter, Robert J. Dybas, repair welding. In: ASM International, AmericanSociety for Metals, Editors, Metals Handbook, Welding, Brazing and Soldering,vol. 6, Metals Park, Ohio; 2002. p. 11037.

[17] Kroes MJ, Watkins WA, Delp F. Aircraft maintenance and repair. 6th ed. NewYork: Glencoe Macmillan/McGraw-Hill; 1993. 648p.

[18] Vega EO, Hallen JM, Villagomez A, Contreras A. Effect of multiple repairs ingirth welds of pipelines on the mechanical properties. Mater Charact2008;59:1498507.UFRJ, Brazil: Federal University of Rio de Janeiro; 1993.[4] Goranson UG. Fatigue issues in aircraft maintenance and repairs. Int J Fatigue

1993;19:S3S21.[5] Payne AO. The fatigue of aircraft structures. Eng Fract Mech 1976;8:157203.[6] Wenner CA, Drury CG. Analysing human error in aircraft ground damage

incidents. Int J Ind Ergonom 2000;26:17799.3. The reverse bending loads signicantly decrease the fatiguestrength of the AISI 4130 steel, whose endurance limit wereabout 37% and 40% of the yielding stress (ry) for specimens1.60 mm and 1.10 mm thick, respectively.

4. The corrosive process strongly reduced the reverse bending fati-gue strength of AISI 4130 steel.

5. In comparison with specimens from the base-material a reduc-tion in bending fatigue strength for TIG welded specimens (OR)was observed. After the rst weld repair (1R), no-subsequentreduction in bending fatigue strength was observed in compar-ison with the original weld. The second weld repair (2R) againresulted in decreasing of the fatigue strength in comparisonwith that subjected to the rst weld repair. Consequently, it isnot recommended or favorable to the ight-safety the secondweld repair (2R) on critical to the ight-safety components.

6. The TIG (re-)welding procedure induced high compressiveresidual stress values in the AISI 4130 steel.

7. It was veried increase of the stress concentration factor, Kt, atthe weld toe with the successive welding repairs, which over-comes the compressive residual stress eld induced by the(re-)welding process. As a result, all the welded fatigue speci-mens fractured at that region.

Acknowledgement

The authors are grateful to FAPESP/Process No. 99/11948-6,CNPq and FUNDUNESP.

References

[1] Flight Safety Foundation Available from ; 2009[accessed 10.2009].

[2] National Civil Aviation Agency Brazil. Available from ; 2004 [accessed 10.2003].

[3] Godefroid LB. Fatigue crack growth under constant and variable amplitude

l Journal of Fatigue 32 (2010) 12001209[19] Liming L, Xin Du, Meili Z, Guoqing C. Research on the microstructure andproperties of weld repairs in TA15 titanium alloy. Mater Sci Eng A 2007;445446:6916.

[20] Shankar K, Weidong Wu. Effect of welding and weld repair on crackpropagation behaviour in aluminium alloy 5083 plates. Mater Des2002;23:2018.

[21] Sun W, Hyde TH, Becker AA, Williams JA. Some key effects on the failureassessment of weld repairs in CrMoV pipelines using continuum damagemodeling. Eng Fail Anal 2005;12:83950.

[22] Dong P, Hong JK, Bouchard PJ. Analysis of residual stresses at weld repairs. Int JPres Ves Pip 2005;82:25869.

[23] Hyde TH, Sun W, Becker AA, Williams JA. Life prediction of repaired welds in apressurised CrMoV pipe with incorporation of initial damage. Int J Pres Ves Pip2004;81:112.

[24] Aloraier AS, Ibrahim RN, Ghojet J. Eliminating post-weld heat treatment inrepair welding by temper bead technique: role bead sequence in metallurgicalchanges. J Mater Process Technol 2004;153154:392400.

[25] Doc IIS/IIW-956-87. (ex. doc. XV-648-87) prepared by commission XVfundamentals of design and fabrication for welding of the IIW,recommendations for repairs and/or strengthening of steel structures.Welding in the world 1988; 26(11/12): 292306.

[26] Ivanov SA, Monim VI, Teodsio JR. New method of X-ray tensometry. In:International conference on experimental mechanics. Lisboa; 1994. p. 75761.

[27] Bannister AC, Ocejo JR, Gutierrez-Solana F. Implications of the yield stress/tensile stress ratio to the SINTAP failure assessment diagrams forhomogeneous materials. Eng Fract Mech 2000;67:54762.

[28] Murakami Y. Inclusion rating by statistics of extreme values and its applicationto fatigue strength prediction and quality control of materials. J Res Nat InstStand Technol 1994;4:99107.

[29] Murakami Y, Usuki H. Quantitative evaluation of effects of non-metallicinclusions on fatigue strength of high strength steels. II: fatigue limitevaluation based on statistics for extreme values of inclusion size. Int JFatigue 1989;5(11):299307.

[30] Makkonen M. Statistical size effect in the fatigue limit of steel. Int J Fatigue2001;23:395402.

[31] Lancaster JF. Metallurgy of welding. 5th ed. London: Chapman and HallIncorporation; 1993. 389p.

[32] Peterson RE. Stress Concentration design factors. 5th ed. New York: JohnWileyand Sons Incorporation; 1966 [155p].

[33] Veidt M. On the Effect of notch radius and local friction on the Mode I andMode II fracture toughness of a high-strength steel. Eng Fract Mech1997;58:22331.

[34] Wei MY, Chen C. Fatigue crack growth characteristics of laser-hardened 4130steel. Scripta Metall Mater 1994;10(31):13938.

[35] Brown TB, Dauda TA, Truman CE, Smith DJ, Memhard D, Pfeiffer W. Predictionsand measurements of residual stress in repair welds in plates. Int J Pres Ves Pip2006;83:80918.

[36] Chiarelli M, Lanciotti A, Sacchi M. Fatigue resistance of MAG welded steelelements. Int J Fatigue 1999;1:1099110.

M.P. Nascimento, H.J.C. Voorwald / International Journal of Fatigue 32 (2010) 12001209 1209

Considerations on corrosion and weld repair effects on the fatigue strength of a steel structure critical to the flight-safetyIntroductionCurrent Welding Repair Procedure and Applications

Material and methodsMaterialWelding and re-welding proceduresCorrosion processReverse bending fatigue testsMicrostructural and microhardness analysesResidual stresses determination

Results and discussionMonotonic testsReverse bending fatigue testsEffect of the corrosion processEffect of the TIG welding/re-welding process

ConclusionsAcknowledgementReferences