A

Da

b

ARRAA

KGVCOVW

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ispciaecsgcp

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Journal of Materials Processing Technology 213 (2013) 1640– 1652

Contents lists available at SciVerse ScienceDirect

Journal of Materials Processing Technology

jou rn al h om epage : www.elsev ier .com/ locate / jmatprotec

study on V-groove GMAW for various welding positions

.W. Choa, S.J. Naa,∗, M.H. Chob, J.S. Leeb

Department of Mechanical Engineering, KAIST, 291 Daehak-ro, Yuseong-gu, Daejeon 305-701, Republic of KoreaTechnical Research Laboratories, POSCO, Pohang, Gyeongsangbuk-Do 790-300, Republic of Korea

a r t i c l e i n f o

rticle history:eceived 15 October 2012eceived in revised form 29 January 2013ccepted 24 February 2013vailable online xxx

a b s t r a c t

This study performed three-dimensional transient numerical simulations using the volume of fluidmethod in a gas metal arc V-groove welding process with and without root gap for flat, overhead, andvertical welding positions. The elliptically symmetric arc models for arc heat flux, electromagnetic forceand arc pressure were used to describe the more accurate molten pool behaviors. The numerical modelsnot only formed a stable weld bead but also simulated the dynamic molten pool behaviors such as over-flow which was not described before. This study analyzed these molten pool flow patterns for various

eywords:as metal arc welding-grooveomputational fluid dynamicsverflowolume of fluid

welding positions and validated the numerical models used by comparing the simulation results withexperimental ones.

© 2013 Elsevier B.V. All rights reserved.

elding position

. Introduction

A circumferential welding process is widely used in shipbuild-ng and construction industries to transport the important fluids,uch as, water and oil. Even though the circumferential weldingrocess is very significant in many industrial fields, installation of aircumferential welding automation system is challenging becauset requires the accurate seam-tracking equipment, precise pipelinelignment, and optimization of welding parameters. Among sev-ral difficulties faced, the optimization of welding parameters inircumferential welding requires an enormous effort to obtain theound and uniform weld beads owing to the various effects ofravity in different welding positions. Therefore, the weld defectsan occur in some welding positions, even with the same weldingarameters being used along the circumferential welding.

With the aim of reducing the weld defects in the welding pro-ess, statistical experimental designs, linear regression modeling,nd neural networks have been used to analyze and optimize theelding process parameters on weld bead geometry. Tay and Bulter

1997) optimized the welding parameters by using neural networksnd the Taguchi method in gas metal arc welding (GMAW) on bead-n-plate (BOP). Kim et al. (2005) predicted the weld bead geometry

y adopting neural networks and regression models for GMA V-roove welding. Cho et al. (2009) predicted the surface profile ofeld beads in flux cored arc welding (FCAW) processes for variable

∗ Corresponding author. Tel.: +82 42 350 3216; fax: +82 42 350 3256.E-mail address: sjoona@kaist.ac.kr (S.J. Na).

924-0136/$ – see front matter © 2013 Elsevier B.V. All rights reserved.ttp://dx.doi.org/10.1016/j.jmatprotec.2013.02.015

welding positions with the Taguchi method and non-linear regres-sion. These works show that the welding process parameters arestrongly correlated to the weld bead geometry. Thus, it is quiteimportant to select the accurate welding process parameters foroptimal weld beads.

Even though these methods could extract the optimal weldingparameters, they do not yield the fundamental reasoning behindthe molten pool formation behavior and the resultant weld beadshape. To overcome this shortcoming, a numerical simulationshould be applied to describe the heat flow patterns and weld beadformations. In most cases, the finite element method (FEM), whichconsidered only conduction heat transfer, was used to predict themolten zone and welding distortions. However, the convection heattransfer was added to predict the molten zone and heat affectedzone (HAZ) in the welding process more accurately. Kim and Na(1989) analyzed the molten pool behaviors such as buoyancy, elec-tromagnetic force (EMF), drag force and surface tension in a gastungsten arc welding (GTAW) process with boundary fitted coordi-nates. In addition, Kim and Na (1992) studied the weld pool surfacedeformation by using the arc pressure in GTAW. Fan et al. (2000)described the molten pool flow patterns in a partially and fullypenetrated weld pool in GTAW with a boundary-fitted coordinate.

The GMAW process is superior to the GTAW process becauseof higher productivity resulted from the molten wire droplet. Kimand Na (1994) performed a three-dimensional (3D) quasi-steady-

state GMAW simulation by the finite difference method (FDM) tocalculate the temperature distribution and convective heat trans-fer with flow patterns. They found that the fingertip molten zonecould be predicted by a convective heat transfer. To calculate the

cessin

swBiKvafawttd

tfnwth(rs

lwa(tabthta

D.W. Cho et al. / Journal of Materials Pro

imulation in detail, the heat and mass transfer of droplets on theeld pool should be considered. Tsao and Wu (1998) and Kim andasu (1998) considered the momentum and enthalpy of droplets

n GMAW and analyzed the droplet effect on weld pool behaviors.im et al. (2003) combined the arc and droplet heat sources as aolumetric heat source model in GMA V-groove welding. Kumarnd Debroy (2007) simulated the molten pool flow in filet GMAWor different welding positions and assumed the welding processs being quasi-steady-state. To satisfy this assumption, however,elding positions in the simulation should be limited and near to

he flat position. Therefore, a quasi-steady-state numerical simula-ion cannot predict the dynamic molten pool flows such as sphericalroplet impingent and unstable weld beads.

The volume of fluid (VOF) simulation technique was used torack the deformation of the weld pool surface due to variable arcorces that were mathematically modeled and implemented in theumerical simulation. This method adopts the transient analysis,hich can detect the free surface variation from the simulation

ime. Therefore, it can predict the unstable weld beads, such asumping and burn-through, as well as stable weld beads. Cao et al.2004) described the droplet impingent on a free surface and theesultant weld bead formation in the GMAW process by commercialoftware (Flow3D).

Additionally, a more complex welding process can also be calcu-ated by VOF. Cho and Na (2006) performed the simulation of laser

elding, which includes very complex physical phenomena suchs multiple reflection and keyhole formation. Moreover, Cho et al.2009) performed the 3D laser-GMA hybrid welding, which adoptshe characteristics of laser welding and GMAW with the followingssumptions: the interaction between the laser and the arc coulde negligible. The VOF method could be also applied to describe

he alloying element distributions and pore generation in the GMAybrid welding process (Cho et al., 2010, 2012). This study analyzedhe molten pool behaviors and weld beads in V-groove GMAW withnd without root gap for various welding positions.

Fig. 1. Schematic sketch of V-groove shapes (a) w

g Technology 213 (2013) 1640– 1652 1641

2. Mathematical formulation

2.1. Material shape and mesh size

This study describes the molten pool dynamics for various weld-ing positions. First, the molten pool analysis without a root gap isperformed, after which the analysis with a 1-mm root gap is per-formed. Fig. 1 shows a schematic sketch of V-groove materials and3D x–y–z coordinate system.

This study used the mesh density as 0.25 mm/mesh. Cho et al.(2013a) found that if the mesh size is larger than 0.25 mm, thevolume of droplet can be lost so that molten pool dynamics aswell as droplet impingent cannot be described accurately. More-over, many previous researches adopt the size of droplet between0.2 mm/mesh and 0.25 mm/mesh in arc welding, laser welding andlaser-arc hybrid welding (Cho and Na, 2009; Cho and Farson, 2007;Cho et al., 2012, 2013a).

2.2. Governing equations

The governing equations used in this study for computationalfluid dynamics (CFD) simulations of a weld pool are the con-tinuity equation, the momentum equation (referred from theNavier–Stokes equation), and the energy equation (Cho and Na,2009; Cho et al., 2010, 2013b). The commercial package Flow-3D isused for the simulation with a VOF equation. The material proper-ties and variables are summarized in Table 1.

- Momentum equation:

�∂ �V + �( �V · ∇) �V = −∇p + ∇2 �V + f (1)

∂t b

- Continuity equation:

∇ · �V = 0 (2)

ithout root gap and (b) with 1-mm root gap.

1642 D.W. Cho et al. / Journal of Materials Processing Technology 213 (2013) 1640– 1652

Table 1Prosperities used in simulation.

Symbol Nomenclature Symbol Nomenclature

k Thermal conductivity Rc Radius of the surface curvature�n Normal vector to free surface � Density (solid: 7.8, liquid: 6.9, g/cm3)� Surface tension Cs Specific heat of solid, 7.26 × 106 erg/g s K�Arc GMAW Arc efficiency in GMAW, 0.56 Cl Specific heat of liquid, 7.32 × 106 erg/g s KV Welding voltage Ts Solidus temperature, 1768 KI Welding current Tl Liquidus temperature, 1798 KJ Current density (I/mm2) Td Droplet temperature, 2400 K�x Effective radius of arc in x-direction, 1.54 mm To Room temperature, 298 K�y Effective radius of arc in y-direction, 0.90 mm hsl Latent heat of fusion, 2.77 × 109 erg/g sfd Droplet frequency (Hz) �d Droplet efficiency in GMAW, 0.24WFR Wire feed rate, 7.5 m/min z1 Vertical distance from top surfacerw Wire diameter in GMAW, 1.2 mm Jze Vertical component of the current densityrd Droplet diameter in GMAW, 1.2 mm Jre Radial component of the current density�0 Permeability of vacuum, 1.26 × 106 H/m B�e Angular component of the magnetic field�m Material permeability, 1.26 × 106 H/m J0 First kind of Bessel function of zero order

-

-

2

h(o

k

tFtVaer

Q

Mcdf00sfl

f

qd Heat input from droplet

PA Arc pressure

VX Velocity of fluid in x-direction, mm/s

Energy equation:

∂h

∂t+ ( �V · ∇)h = 1

�∇ · (∇T), where h = CpT + f (T)Lf (3)

f (T) =

⎡⎢⎢⎣

0 (T ≤ Ts)

T − Ts

Tl − Ts(Ts < T < Tl)

1 (Tl ≤ T)

(4)

VOF equation:

dF

dt= ∂F

∂t+ ( �V · ∇)F = 0 (5)

.3. Boundary conditions

The energy on the top free surface is balanced among the arceat flux (QA), heat dissipation by convection (Qconv) and radiationQrad), and heat loss due to evaporation (Qevap). The energy balancen the top surface is expressed as the following equation:

∂T

∂�n = QA − Qconv − Qrad − Qevap. (6)

Many previous researches applied axisymmetric Gaussian dis-ributed arc models (Cao et al., 2004; Cho and Na, 2009; Cho andarson, 2007; Cho et al., 2010). However, Cho et al. (2013b) usedhe elliptically symmetric arc heat flux and arc pressure models in-groove GMA welding and found that the models were valid forpplication to the simulation. Therefore, this study also applies thelliptically symmetric model that contains two different effectiveadii (�x = 1.50 mm, �y = 0.90 mm) of the arc plasma in Eq. (7).

A(x, y) = �Arc GMAW VI

2�x�yexp

(−

(x2

2�2x

)−

(y2

2�2y

))(7)

Previous researchers such as Cho and Na (2009), Dupont andarder (1995), and Cho et al. (2013b), used a total GMAW arc effi-

iency of 0.8 which includes the heat transfer by arc and moltenroplets; moreover, the heat input efficiency of the droplets wasound to be 0.24 from Eqs. (8)–(11). Therefore, an arc efficiency of.56 is used in this study, together with the droplet efficiency of.24 by the droplet transfer. The energy balance on the bottom freeurface is expressed in a manner similar to Eq. (7), but the arc heat

ux model should be ignored.

d = 3r2wWFR

4r3d

, (8)

J1 First kind of Bessel function of first orderCy Thickness of the workpiece

qd = 43

r3d � [Cs(Ts − To) + Cl(Td − Ts) + hsl] fd, (9)

�d = qd

VI, (10)

�Arc GMAW = 0.8 − �d. (11)

For the pressure boundary conditions, the following Eq. (12) isused at the free surface.

p = pA + �

Rc, (12)

pA = �0IJ

4, (13)

pA(x, y) = �0I2

42�x�yexp

(−

(x2

2�2x

)−

(y2

2�2y

)), (14)

Lin and Eagar (1986) found that the current and current den-sity are linearly proportional to the arc pressure (pA) in Eq. (13).Therefore, this study assumes that the distribution of the arc pres-sure follows the distribution of the current density. The ellipticallysymmetric arc plasma pressure can therefore be modeled as shownin Eq. (14).

2.4. Electromagnetic force

Cho et al. (2013b) calculated the EMF distribution by mappingcoordinates on V-groove welding processes. They concluded thatthe coordinates mapping of EMF is necessary in GTAW; however, itnot useful in GMAW. The axisymmetric EMF model must be mod-ified because the arc heat flux model is elliptically symmetric inV-groove welding. Eqs. (15)–(17) can be used to formulate Eq. (18),which is a simple elliptically symmetric model that contains aneffective radius of the welding arc for the x-direction as well as theelliptical radius (re). This model modifies the current density andelectromagnetic field and ultimately determines the EMF for the x,y, and z directions (Fx, Fy, Fz).

x2

a2+ y2

b2= 1, (15)

k1 = b = �y, (16)

a �x

x2 + y2

k21

= r2e , (17)

D.W. Cho et al. / Journal of Materials Processing Technology 213 (2013) 1640– 1652 1643

Table 2Common welding conditions (with or without root gap).

Variable Value

Wire feed rate 7.5 m/minVoltage 25 VElectrode YGW15, ̊ = 1.2 mmCurrent 255 AWelding speed 10 mm/s (cases 1–7), 20 mm/s (case 8)

e

J

J

B

F

F

F

2

(a

3

nCGTwfl

Table 3Welding positions (without root gap).

Welding position

Case 1 Flat

CTWD 20 mmTorch angle 90◦

Shielding gas 80%Ar–20%CO2, 20 l/min

xp

(− x2

2�2x

− y2

2�2y

)= exp

(− r2

e

2�2x

), (18)

ze = I

2

∫ ∞

0

�J0(�re) exp

(−�2�2

x

12

)sinh[�(cy − z1)]

sinh(�cy)d�, (19)

re = I

2

∫ ∞

0

�J1(�re) exp

(−�2�2

x

12

)cosh[�(cy − z1)]

sinh(�cy)d�, (20)

�e = �mI

2

∫ ∞

0

J1(�re) exp

(−�2�2

x

12

)sinh[�(cy − z1)]

sinh(�cy)d�, (21)

x = −JzeB�ex

re, (22)

y = −JzeB�ey

re, (23)

z = JzeB�e. (24)

.5. Other welding models

In this study, the same models as those used in previous studiesCho and Na, 2009; Cho et al., 2010, 2013b), such as buoyancy forcend drag force of arc plasma are used.

. Results and discussion

In this paper, all the simulation were conducted by transientumerical analysis for 3 s after beginning except case 8 in Table 5.ho et al. (2013b) performed the numerical simulation of V-groove

MAW in a flat position under the following conditions (seeable 2). This study uses the same welding conditions with varyingelding positions and analyzes not only the bead formation in theat position but also that in different welding positions.

Fig. 2. Simulation result in case 1 (a) 3D bead shape and

Case 2 OverheadCase 3 Vertical-up

3.1. Without root gap

Table 3 lists the different welding positions used in the simu-lation without the root gap. In case 1, a uniform and incompletelypenetrated weld bead can be formed for a three-dimensional weldbead, a transverse cross section, and a transverse cross section atthe center line, as shown in Fig. 2.

Fig. 3 describes the stable molten pool flow patterns on a longi-tudinal cross section where the maximum velocity of flow towardthe rear welding direction is 210 mm/s, which is not enough tomake a weld bead with complete penetration. In addition, thesloped surface of the V-groove melts earlier than the weld seamowing to the geometrical shape of the V-groove as shown in Fig. 4.These molten pool flow patterns also disturb a fully penetratedweld bead, even though EMF and arc pressure from arc plasmatend to penetrate the weld pool downward. Finally, the moltenpool flows can be circulated on the V-groove slope and, in turn,can induce the incomplete penetration in case 1.

In case 2 (overhead position), uniform and stable molten poolflow patterns are described, although the reverse gravity vectoraffects the molten pool (see Figs. 5–7). These flow patterns are verysimilar to case 1; moreover, they form the incompletely penetratedweld bead in case 2. By comparing the case 1 with case 2, it is possi-ble to understand that a gravity vector in the z direction has only aweak effect on the molten pool and the final bead formation. Fig. 6describes the stable molten pool flow patterns on a longitudinalcross section where the maximum velocity of flow toward the rearwelding direction is 215 mm/s, which is insufficient to make a fullypenetrated weld bead in case 2.

Looking at the cross section of molten pool flow patterns showsthe similarities in case 1 and case 2, where both show the forma-tion of a incompletely penetrated weld bead (see Fig. 7). However,bead widths and reinforcement heights between case 1 and case2 are different owing to the direction of gravity force, as shownin Table 4. In case 2, the gravity force draws the molten pool tothe top of weld bead, causing a narrower bead width and higherreinforcement than that formed in case 1.

In the vertical position, the welding direction is the same asor opposite to the direction of gravity according to the vertical-down or vertical-up position respectively, so that the molten

(b) experiment and simulation (Cho et al., 2013b).

1644 D.W. Cho et al. / Journal of Materials Processing Technology 213 (2013) 1640– 1652

Fig. 3. Calculated temperature profiles and flow patterns on a longitudinal cross-section in case 1.

Fig. 4. Calculated temperature profiles and flow patterns on a transverse cross-sectional (x = 3.0 cm) in case 1.

Fig. 5. Simulation result in case 2 (a) 3D bead shape and (b) experiment and simulation.

Fig. 6. Calculated temperature profiles and flow pat

terns on a longitudinal cross-section in case 2.

D.W. Cho et al. / Journal of Materials Processing Technology 213 (2013) 1640– 1652 1645

Fig. 7. Calculated temperature profiles and flow patterns on a transverse cross-section (x = 3.0 cm) in case 2.

Table 4Comparision of simulation result with experiment.

Experiment Simulation result

Reinforcement hight (mm) Bead width (mm) Reinforcement hight (mm) Bead width (mm)

Case 1 0.86 5.53 0.8 5.28Case 2 0.97 5.13 0.95 5.02

exper

porasflwtlFc

Fig. 8. Simulation and

ool behavior can be very sensitive and dynamic compared tother welding positions. This study adopts the vertical-up position,eferred to as case 3, which induces the dynamic weld beads suchs humping on the top surface and melt-through on the bottomurface, as shown in Fig. 8. In case 1 and case 2, the molten poolows can be circulated in a clockwise direction to form the stableeld beads as shown in Figs. 3 and 6, respectively. On the contrary,

he molten pool can be accelerated to the rear welding direction inine with the direction of gravity in the case of vertical-up position.ig. 9(a) describes the unstable flow patterns in the longitudinalross section where the maximum velocity of molten pool flow

Fig. 9. Calculated temperature profiles and flow pat

iment result in case 3.

toward the rear welding direction is 390 mm/s, which results inthe partial circulation and the formation of a humping bead on thetop surface. Thus, a more active convection heat transfer can meltthe V-groove material and form a fully penetrated weld bead. Thedashed circle region in Fig. 9(b) solidifies earlier because the heatis dissipated by a conduction transfer. This solidified region breaksthe molten pool channel and heads to the top surface; therefore,

the rest of the molten pool flows to the bottom surface and formsa melt-through bead. The numerical models used in the simula-tion can be validated by comparing the simulation results with theexperimental ones. However, the welding conditions presented in

terns on a longitudinal cross-section in case 3.

1646 D.W. Cho et al. / Journal of Materials Processing Technology 213 (2013) 1640– 1652

Fig. 10. Simulation result of fusion zone on a longitudinal cross-section in case 3.

Table 5Welding positions (1 mm root gap).

Position (torch angle)

Case 4 Flat (90◦)Case 5 Flat (45◦ , front direction)Case 6 Overhead (90◦)

Toavow

GftmWacHbtp

3

brca

ltsmtts

c

welding process. Fig. 16 compares the simulation results of the weld

Case 7 Vertical downward (90◦)Case 8 Vertical downward (90◦), welding speed (20 mm/s)

able 2 are not suitable for applying the V-groove GMAW with-ut a root gap because these conditions result in weld defects suchs incomplete penetration, humping and melt-through beads forarious welding positions (Fig. 10). Resultantly, it is necessary toptimize the welding parameters in V-groove GMAW for variouselding positions.

Several studies focused humping bead formation in high speedMAW. Cho and Farson (2007) calculated the humping bead

ormation in high-speed GMAW by VOF method. They clarifiedhe physical mechanisms of the humping phenomenon from the

olten pool fluid patterns and molten pool solidification. Chen andu (2010) observed molten pool behaviors of humping bead with

high speed camera in high-speed GMAW. Those studies wereonducted in high speed welding only for flat welding position.owever, this study also obtained humping bead and melt throughead not in flat position but in vertical upward positions, therefore,he mechanism to make the humping bead is also different fromrevious studies.

.2. With 1-mm root gap

In order to prevent the weld defects and obtain a sound weldead, the V-groove joint between the plates is opened with a 1-mmoot gap. The welding conditions are the same as those in previousases (see Table 2); however, the torch angles and welding positionsre varied, as given in Table 5.

In case 4, droplets from the molten wire impinge perpendicu-arly on the weld pool surface. Under this condition, the surfaceension on the bottom surface hardly sustains the variable forcesuch as droplet impingement, EMF, and arc pressure. Thus, theolten pool can leak from the bottom surface where a burn-

hrough bead is formed, as shown in Fig. 11. However, Fig. 12 showshat the stable weld beads can be formed on the top and bottom

urfaces as in case 5.

Droplets are added to the weld pool so that the molten fluidsan accumulate completely (see Fig. 13(a)). Next, droplets slantly

Fig. 12. Stable weld

Fig. 11. Burn-through bead in case 4 (top surface).

impact and push the small amount of weld pool to penetrate intothe root gap. As shown in Fig. 13(b), the penetrated molten poolsolidified earlier owing to conduction heat transfer. When the cen-ter of the arc plasma reaches the pre-solidified region (see Fig. 13(c)and (d)), it can be melted again owing to EMF and arc pressuresuppressing the molten pool downward, making a fully penetratedweld bead. Meanwhile, the molten fluid flows to the rear anddownward direction because the pre-solidified region reflected themolten pool. Thus, the downward momentum of the molten poolbecame smaller than in case 4, so the fully penetrated molten poolcould not leak from the bottom surface and the surface tensioncould sustain the variable arc forces in case 5. Finally, this condi-tion resulted in the formation of a uniform, fully penetrated, andsound weld bead.

Fig. 14(a)–(f) shows the temperature profiles and the moltenpool flow patterns on a transverse cross-section in case 5. The over-flowed molten pool penetrates into the root gap (see Fig. 14(a))and then solidifies as shown in Fig. 14(b). Next, this pre-solidifiedregion melts again (Fig. 14(c)) and forms a fully penetrated weldbead (Fig. 14(d)). At the bottom surface, the molten pool existswhile the upper part is solidified by conduction heat transfer;therefore, the gravity pulls the molten metal downward to resultin a higher back-bead height (Fig. 14(e)). Finally, a fully pene-trated molten pool spreads outside owing to the surface tension(Fig. 14(f)).

The molten pool overflow phenomenon can be observed by ahigh-speed camera as shown in Fig. 15. Just before overflow, themolten pool has fully accumulated in the V-groove and the arcplasma is located in a higher position. After overflow, however,the molten pool penetrates in the root gap; thus, the arc plasmais located in a lower position. If there is no overflow pattern in V-groove welding, however, the arc position can be stable during the

bead in case 5.

bead cross section with the experimental results to verify the weld-ing models and algorithms used in this work. Until now, there wasno research which described the molten pool overflow behavior.

D.W. Cho et al. / Journal of Materials Processing Technology 213 (2013) 1640– 1652 1647

Fig. 13. Calculated temperature profiles and flow patterns on a longitudinal cross-section in case 5.

tterns

Tb

ss

Fig. 14. Calculated temperature profiles and flow pa

his paper firstly observed and described the molten pool overflow

ehavior in welding simulation area.

In case 6, the molten pool overflow pattern is observed by CFDimulation, which is similar to case 5. The penetrated molten poololidifies in the root gap, and then melts again by the arc heat source

on a transverse cross-section (x = 2.7 cm) in case 5.

and arc forces to form a stable bead. In overhead welding, the direc-

tion of gravity is opposite to the direction of flat position welding;therefore, different molten pool flow patterns and the resultantbead shapes can be formed. Fig. 17(a) and (b) shows the temper-ature profiles and the molten pool flow patterns on a longitudinal

1648 D.W. Cho et al. / Journal of Materials Processing Technology 213 (2013) 1640– 1652

Fig. 15. Molten pool overflow from a high speed camera.

ctpsm

Fig. 16. Fusion zone profiles (case 5).

ross-section in case 6. The dashed region is far from the arc cen-

er, so arc forces such as EMF and arc pressure prevent the moltenool from flowing downward. On the contrary, the molten pool istretched at the upper part owing to the gravity force, so the dashedolten region can remain as a molten fluid for a longer time.

Fig. 17. Calculated temperature profiles and flow pat

Fig. 18. Calculated temperature profiles and flow pattern

Fig. 19. Fusion zone profiles in case 6.

Fig. 18 shows the temperature profiles and the molten poolflow patterns on a transverse cross-section in case 6. After themolten pool overflows, a small amount of molten fluid pene-trates into the root gap and solidifies as shown in Fig. 18(a). Thisregion can be melted again when the arc center reaches it, andforms a fully penetrated weld bead as shown in Fig. 18(b). Eventhough the molten pool fully penetrates the root gap, it is difficultto achieve the sufficient back-bead height, because gravity forceextracts the molten pool upward while surface tension pulls thefluid to the outside on the bottom surface. However, a convex topweld bead can be formed owing to the gravity force, as shown inFig. 18(c). The splashed molten pool during welding process can beattached in the bottom surface; however, it hardly affects the over-all weld bead shape because the amount of splashed molten poolis very tiny. The simulation model can be validated by compar-ing the simulation result with that of the experiment, as shown inFig. 19.

Weld defects such as humping and melt-through beading are

clearly observed in a vertical-upward position in case 3 owing to thegravity force, as mentioned above. To prevent these weld defects,the vertical-downward position is used, as given in Table 4, with

terns on a longitudinal cross-section in case 6.

s on a transverse cross-section (x = 2.6 cm) in case 6.

D.W. Cho et al. / Journal of Materials Processing Technology 213 (2013) 1640– 1652 1649

Fig. 20. Calculated temperature profiles and flow patterns on a longitudinal cross-section in case 7.

h pro

asoFittovtrbts

mrbo

Fig. 21. Detailed flow patterns whic

1-mm root gap. In case 7, the direction of gravity force is theame as the welding direction, so that it can induce a molten poolverflow pattern in a vertical-down welding position as shown inig. 20(a)–(d). The penetrated molten pool in the root gap is solid-fied early (Fig. 20(c)); however, it cannot be melted again evenhough the arc center reaches the solidified region. This is due tohe size of the circulating molten pool being smaller than that inther cases because of the direction of gravity force. The maximumelocity of the molten pool flow toward the rear welding direc-ion is 130 mm/s, which is insufficient to melt the penetrated solidegion again. Moreover, case 7 cannot induce a uniform weld beadut results in the formation of two weld defects: lack of penetra-ion (A) and lack of fusion (B) on the longitudinal cross-section, ashown in Fig. 20(d).

Figs. 21(a) and (b) and 22 show the cross-sectional views of

olten pool flow patterns that induce a lack of penetration. The

eason for this lack of penetration is that the solid region in the redoxed area reflects the molten fluid, so that it goes upward with-ut penetrating (Figs. 21(a) and 22(a)). Although the arc forces are

duce a lack of penetration in case 7.

acting on the region to make the molten pool flow downward, asshown in Fig. 22(b), the forces are insufficient to cause the pene-tration into the root gap and result in a incompletely penetratedregion.

Fig. 23 shows the molten pool flow patterns on a transversecross-section for the lack of fusion. The penetrated molten poolbarely melts the root face and then solidifies quickly. Additionally,this part cannot be melted again as mentioned above; therefore, awelding defect, such as lack of fusion, can occur. Fig. 24 comparesthe simulation results with the experimental results for two welddefect regions, with the aim of validating the models used in thesimulation. By observing both the top and bottom surface of weld-ing experimental specimen, it is also possible to conclude that thelack of penetration and the lack of fusion regions can coexist underthe same welding condition of the case 7, as shown in Fig. 25.

To prevent the formation of weld defects in the vertical-downward position, it is necessary to change the weldingparameters to avoid the molten pool overflow. One of the rea-sons for generating the overflow pattern in a vertical-downward

1650 D.W. Cho et al. / Journal of Materials Processing Technology 213 (2013) 1640– 1652

erns w

pa(snhcm4p

Fig. 22. Calculated temperature profiles and flow patt

osition is the accumulation of the molten pool. Thus, this studydopted a welding speed twice (20 mm/s) that in other cases1–7) to reduce the accumulated molten pool. Simulation resultshow that overflow patterns are not observed on the longitudi-al and transverse cross-sections, as shown in Figs. 26 and 27. Theigher welding speed causes a greater molten pool circulation as

an be observed on the longitudinal cross-section. Moreover, theaximum velocity of flow toward the rear welding direction is

00 mm/s, which is enough to melt the root face and form a fullyenetrated weld bead. Finally, a uniform and stable weld bead can

Fig. 23. Calculated temperature profiles and flow pattern

Fig. 24. Fusion zone profiles from expe

hich produce a lack of penetration (part A) in case 7.

be formed in the vertical-down position with a 1-mm root gap.Fig. 28 compares the simulation result with the experimental onewith the aim of validating the models used in the simulation forhigher welding speed in the vertical-down position.

All the simulations were conducted by transient numericalanalysis which described dynamic molten pool behavior such as

humping and overflow. Even though it is also meaningful to com-pare the temperature contours at the same time instant for severalcases, it is more effective to show the dynamic molten pool forma-tion separately.

s which produce a lack of fusion (part B) in case 7.

riment and simulation in case 7.

D.W. Cho et al. / Journal of Materials Processing Technology 213 (2013) 1640– 1652 1651

Fig. 25. Weld bead surfaces in case 7.

Fig. 26. Calculated temperature profiles and flow patterns from a cross-sectional side view in case 8.

Fig. 27. Calculated temperature profiles and flow patterns on a transverse cross-section (x = 2.6 cm) in case 8.

Fig. 28. Fusion zone profiles from experiment and simulation in case 8.

1 cessin

4

sis

(

itflcmigp

A

KtK

R

C

C

652 D.W. Cho et al. / Journal of Materials Pro

. Conclusions

This study demonstrated the dynamic molten pool behaviorsuch as humping, melt-through, and overflow for various weld-ng positions in V-groove GMAW. The results of this work can beummarized as follows:

(a) Without the root gap, it is difficult to form a fully penetratedweld bead in the flat and overhead positions, while humpingand melt-through beads are formed in the vertical-upwardposition under the same welding condition.

b) With a 1-mm root gap, the molten pool overflow patterns canbe described for various welding positions under the givenwelding conditions. The overflow patterns in some weldingpositions do not induce the weld defects, while a weld beadwith incomplete penetration can be formed in the vertical-downward position. Thus, it is necessary to avoid the overflowpatterns in such a case by increasing the welding speed.

Previous studies could obtain weld defects such as hump-ng and lack of penetration by numerical simulation. However,hey performed welding simulations and experiments only in theat position. This study performed the positional welding pro-ess in V-groove GMAW by transient analysis. Therefore, dynamicolten pool behaviors such as humping in vertical upward weld-

ng positions and overflow molten pool behavior with a 1 mm rootap in various welding positions were firstly introduced in thisaper.

cknowledgements

The authors gratefully acknowledge the support of the Brainorea 21 project, POSCO and the grant (No. 2010-0027749) from

he National Research Foundation of Korea, which is funded by theorean Ministry of Education, Science and Technology.

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