modelling of angular distortion of double-pass butt-welded plate

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Modelling of angular distortion of double-pass butt-welded plate M M Mahapatra 1 , G L Datta 1 *, B Pradhan 1 , and N R Mandal 2 1 Department of Mechanical Engineering, Indian Institute of Technology, Kharagpur, India 2 Department of Ocean Engineering and Naval Architecture, Indian Institute of Technology, Kharagpur, India The manuscript was received on 31 August 2007 and was accepted after revision for publication on 6 November 2007. DOI: 10.1243/09544054JEM995 Abstract: Adequate top and bottom side weld reinforcements are important for single-side submerged arc welded butt joints as the shrinkage forces generated due to the solidification of top side weld reinforcements and bottom side weld reinforcements can cancel out each other to minimize angular distortions. In the present investigation it was experimentally established during submerged arc welding (SAW) that by using proper process parameters adequate bottom and required top weld reinforcements could be obtained by using two welding passes for 12 mm thick square butt joints. During the experimental investigation the first pass of weld- ing was used for adequate bottom reinforcement and for filling up the root gap. The second welding pass was used to achieve the required top reinforcement. The first pass weld acted like a restraint to angular distortion induced due to the weld of the second pass. The process was modelled using three-dimensional finite element analysis considering a distributed moving heat source, reinforcements, filler material deposition in each pass of welding, and temperature- dependent material properties. Keywords: square butt joints, double-pass submerged arc welding, finite element analysis, transient thermal analysis, element birth and death method, thermomechanical analysis and angular distortions 1 INTRODUCTION Although adequate top and bottom weld reinforce- ments are desired in a single pass of welding for mini- mizing angular distortions, it is not always possible and a second pass is often used to attain the top reinforce- ment. In heavy structural welding works bead reinfor- cements (i.e. bead width and height) are normally prescribed for better structural integrity. These reinfor- cements can be achieved by using automatic processes like submerged arc welding (SAW) with proper process parameters. If the first weld pass produces adequate bottom reinforcement then the resulting angular dis- tortion after the second weld pass will be less because the bottom reinforcement, resulting from the first pass, would act like a constraint in the second pass. Although the SAW process is used widely in heavy structural fabrications, works related to the prediction of angular distortions of multipass submerged arc welded square butt joints (i.e. joints without any edge preparation) with top and bottom weld reinforcements using three-dimensional finite element analysis are rarely found in the published literature. There is there- fore a need to develop an experimental procedure and three-dimensional model for achieving and predicting angular distortions caused by multipass submerged arc welded butt welding with top and bottom rein- forcements. In the present work single-side double- pass submerged arc welded square butt welding of 12 mm thick mild steel plates as carried out incorporat- ing an adequate bottom weld reinforcement in the first pass and required top weld reinforcement in the second pass. The process was modelled using three- dimensional finite element analysis. 2 LITERATURE REVIEW In earlier days of welding simulations the tempera- ture distributions, residual stresses, and distortions in single-pass welded joints were used in predictions by considering a two-dimensional approximation of a three-dimensional problem [1, 2]. Free and Porter Goff [3] had used a two-dimensional plane strain *Corresponding author: Department of Mechanical Engineer- ing, Indian Institute of Technology, Kharagpur, West Bengal 721 302, India. email: [email protected] 391 JEM995 Ó IMechE 2008 Proc. IMechE Vol. 222 Part B: J. Engineering Manufacture

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Page 1: Modelling of Angular Distortion of Double-pass Butt-welded Plate

Modelling of angular distortion ofdouble-pass butt-welded plateM M Mahapatra1, G L Datta1*, B Pradhan1, and N R Mandal2

1Department of Mechanical Engineering, Indian Institute of Technology, Kharagpur, India2Department of Ocean Engineering and Naval Architecture, Indian Institute of Technology, Kharagpur, India

The manuscript was received on 31 August 2007 and was accepted after revision for publication on 6 November 2007.

DOI: 10.1243/09544054JEM995

Abstract: Adequate top and bottom side weld reinforcements are important for single-sidesubmerged arc welded butt joints as the shrinkage forces generated due to the solidificationof top side weld reinforcements and bottom side weld reinforcements can cancel out each otherto minimize angular distortions. In the present investigation it was experimentally establishedduring submerged arc welding (SAW) that by using proper process parameters adequatebottom and required top weld reinforcements could be obtained by using two welding passesfor 12mm thick square butt joints. During the experimental investigation the first pass of weld-ing was used for adequate bottom reinforcement and for filling up the root gap. The secondwelding pass was used to achieve the required top reinforcement. The first pass weld actedlike a restraint to angular distortion induced due to the weld of the second pass. The processwas modelled using three-dimensional finite element analysis considering a distributed movingheat source, reinforcements, filler material deposition in each pass of welding, and temperature-dependent material properties.

Keywords: square butt joints, double-pass submerged arc welding, finite element analysis,transient thermal analysis, element birth and death method, thermomechanical analysis andangular distortions

1 INTRODUCTION

Although adequate top and bottom weld reinforce-ments are desired in a single pass of welding for mini-mizing angular distortions, it is not always possible anda second pass is often used to attain the top reinforce-ment. In heavy structural welding works bead reinfor-cements (i.e. bead width and height) are normallyprescribed for better structural integrity. These reinfor-cements can be achieved by using automatic processeslike submerged arc welding (SAW) with proper processparameters. If the first weld pass produces adequatebottom reinforcement then the resulting angular dis-tortion after the second weld pass will be less becausethe bottom reinforcement, resulting from the firstpass, would act like a constraint in the second pass.

Although the SAW process is used widely in heavystructural fabrications, works related to the predictionof angular distortions of multipass submerged arcwelded square butt joints (i.e. joints without any edge

preparation) with top and bottom weld reinforcementsusing three-dimensional finite element analysis arerarely found in the published literature. There is there-fore a need to develop an experimental procedure andthree-dimensional model for achieving and predictingangular distortions caused by multipass submergedarc welded butt welding with top and bottom rein-forcements. In the present work single-side double-pass submerged arc welded square butt welding of12mm thickmild steel plates as carried out incorporat-ing an adequate bottom weld reinforcement in the firstpass and required top weld reinforcement in thesecond pass. The process was modelled using three-dimensional finite element analysis.

2 LITERATURE REVIEW

In earlier days of welding simulations the tempera-ture distributions, residual stresses, and distortionsin single-pass welded joints were used in predictionsby considering a two-dimensional approximation ofa three-dimensional problem [1, 2]. Free and PorterGoff [3] had used a two-dimensional plane strain

*Corresponding author: Department of Mechanical Engineer-

ing, Indian Institute of Technology, Kharagpur, West Bengal

721 302, India. email: [email protected]

391

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analysis for residual stress prediction in weldedplates, but it resulted in higher stress predictions asthe thermal expansion in the longitudinal directioncould not be accommodated in their model.Michaleris [4] had predicted residual stresses in mul-tipass girth welds of thin and thick pipes using quasi-static stress analysis. Michaleris and DeBiccari [5]combined welding simulations with three-dimen-sional structural analyses in a decoupled approachto evaluate welding-induced buckling in panel struc-tures. They had used a kinematic work hardeningmaterial model for simulating the plastic behaviourof mild steel [5]. Lindgren [6] discussed complexitiesinvolved in simulation of welding processes andpresented improved material models for better pre-diction of residual stresses and distortions for three-dimensional modelling. Teng et al. [7] had usedtemperature-dependent material properties andconvection cooling in their two-dimensional finiteelement model for predicting residual stresses anddistortions in T-joint fillet welds. Park et al. [8] hadstudied the effects of mechanical constraints onangular distortion of welded joints. The effects of pro-cess parameters on certain responses in welding ofcurved mild steel plates were investigated by Awang[9]. Tsirkas et al. [10] evaluated distortions in laser-welded shipbuilding parts using a local–global finiteelement approach. They had obtained the residualplastic strains and stiffness from the local model andutilized the same in the global analysis for predictingthe distortion of the whole part [10]. Pathak and Datta[11] predicted the temperature distribution andmicrostructure zones of submerged arc-welded jointsusing the distributed circular spread of arc in theirthree-dimensional finite element models. Fanous etal. [12] used three-dimensional finite element model-ling using element birth and element movementmethods for predicting temperature distributionsand residual stresses with the help of the distributedcircular spread of arc in the thermal model for pre-dicting the temperature distributions. Fassani andTrevisan [13] had presented an analytical solution topredict temperature fields in multipass welding usingdistributed Gaussian heat sources and compared theresults with those of a concentrated heat source;they opined that the thermal profiles obtained fromthe distributed heat source models were more reliablethan those obtained from the concentrated heatsource models. Tsai and Jung [14] modelled the angu-lar distortion of T-joints using plasticity-based distor-tion analysis. Dhingra andMurphy [15] did numericalanalyses of several joints including fillet joints for pre-dicting welding-induced distortions in thin-walledstructures. Jiang et al. [16] used a three-dimensionalaxis-symmetric thermomechanical finite elementmodel to simulate the multipass welding processusing the established data [17] for predicting residual

stresses. They opined that the predictions from thethree-dimensional model were better when com-pared with the experimental data than those fromthe two-dimensional models [16].

From the literature review for single- and multipasswelding the need and suitability of three-dimensionalthermomechanical finite element modelling with amoving distributed heat source was realized and car-ried out in the present investigation for predicting thetemperature distributions and angular distortions ofdouble-pass submerged arc-welded butt welds.

3 MODELLING METHODOLOGY

In the present work three-dimensional finite elementanalyses were carried out for predicting the transienttemperature distributions and angular distortions ofdouble-pass submerged arc-welded square buttjoints (i.e. joints without any edge preparation) bytaking into consideration the following:

(a) moving heat source;(b) temperature-dependent thermal and mechanical

material properties;(c) incorporating the joint geometry including top

and bottom reinforcements into the modelling;(d) layer-wise application of heat flux for each pass

of welding;(e) element activation and deactivation for incorpor-

ating filler material deposition in each pass;(f) deactivation of elements of the second weld pass

while applying the heat flux and subsequentcooling to the elements of the first pass;

(g) incorporating an appropriate material model forsimulating elastic–plastic behaviour of the mildsteel weld and base metal.

Heat flux was applied layer-wise for each pass ofwelding, while considering filler material depositionand moving heat source to obtain the transient ther-mal profiles. The thermal profiles were matched withthe experimental data. Angular distortions before andafter welding of the plates were also measured. Tran-sient thermal and non-linear structural analyses werecarried out for predicting angular distortion. Themodel was further verified by comparing the pre-dicted and experimentally obtained angular distor-tions. The composition of the mild steel plates usedin the experiments is shown in Table 1. The thermo-mechanical properties of mild steel [2, 9, 11] usedfor modelling temperature distributions and distor-tions are shown in Table 2. Solidus (Tsolidus) and liqui-dus (Tliquidus) temperatures of the mild steel used inthe analysis were considered to be 1435 and 1500 �Crespectively [11]. The density of the mild steel usedin the analysis was taken as 7850 kg/m3.

392 M M Mahapatra, G L Datta, B Pradhan, and N R Mandal

Proc. IMechE Vol. 222 Part B: J. Engineering Manufacture JEM995 � IMechE 2008

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3.1 Thermal model

The heat source was modelled as a distributed heatflux depending on arc spread for each pass of welding.The rate of arc travel, current, and voltage were variedand these parameters were noted along with the tem-perature data. During experiments the top and bottomside temperatures for each pass of welding away fromthe weld line were measured using thermocouples.Due to the removal procedure of slag and flux granulesjust after the first pass welding the topside thermocou-ples sometimes got damaged and readings becameerroneous. The topside thermocouples were thereforenot useful for recording the temperature histories forboth welding passes. The bottom side thermocoupleswere not affected during the slag removal procedure,so the temperature histories for both the passes couldbe measured from the bottom side thermocouplesonly; the same were used for validation of the resultsobtained from the thermal model. It is not possible tomeasure the spread of arc in the SAW process becausethe arc is covered with flux granules. However, theradius of arc spread of each weld pass was estimated[11] by considering the electrode diameter and beadreinforcements of welds formed during experiments.These arc radii were used for transient thermal analysisof the moving arc for each pass of welding and thetemperature profiles were verified with the experimen-tally measured ones. The moving heat load applied inthe finite element model was taken as a distributedheat flux, as given by

qsupðrÞ ¼ 3Qe

pr2exp �3

r

r

� �2� �

ð1Þ

where r is the region in which 95 per cent of heat fluxis concentrated [11, 12, 18, 19] and

Qe ¼ hVIwhere

Qe ¼ arc power (W)h ¼ arc energy transfer efficiencyV ¼ arc voltage (V)I ¼ arc current (A)

A schematic diagram of a square butt joint (a buttjoint without any edge preparation) is shown inFig. 1. Meshing and modelling of the submergedarc-welded double-pass square butt joints with topand bottom reinforcements are shown in Fig. 2. Thethermal model of the present investigation is basedon the equations of heat conduction. During theSAW the weld zone is covered with flux granules.Therefore, convection loss is not assumed for theweld zone. While heat is conducted throughout themodel due to the heat input from the arc, it isassumed that surface exposed to the atmosphereother than the weld zone is also subjected to theloss of heat by convection. In the present work theheat flux in the arc spread was assumed to follow aGaussian distribution for predicting the temperaturedistribution in the weldment. By using the distribu-ted heat source the temperature at the middle of theweld line can be easily predicted. Many authors[11–18] had also assumed a ‘distributed heat source’for weld thermal modelling for predicting tempera-ture distribution. Moreover, the plate sizes wereenough because for the finite element method(FEM) analysis the weld area was further dividedinto many smaller areas and as the heat source

Table 2 Thermal and mechanical properties of mild steel [2, 9, 11]

Temperature(�C)

Thermalconductivity(W/m K)

Specific heat(J/kg K)

Enthalpy(J/m3) Poison’s ratio

Yield stress(MPa)

Young’smodulus(GPa)

Thermalexpansioncoefficient(10�6/�C)

0 51.9 450 1 ·109 0.2786 290 200 10100 51.1 499.2 2 ·109 0.3095 260 200 11300 46.1 565.5 2.65· 109 0.331 200 200 12450 41.05 630.5 3.8·109 0.338 150 150 13550 37.5 705.5 4.1·109 0.3575 120 110 14600 35.6 773.3 4.55· 109 0.3738 110 88 14720 30.64 1080.4 5 ·109 0.3738 9.8 20 14800 26 931 5.23· 109 0.4238 9.8 20 151450 29.45 437.93 9 ·109 0.4738 — 2 —1510 29.7 400 1.1·1010 0.499 — 0.2 —1580 29.7 735.25 1.1·1010 0.499 0.0098 0.00002 —5000 42.2 400 1.25· 1010 0.499 0.0098 0.00002 15.5

Table 1 Composition of the steel used in the experiments

C (%) Si (%) Mn (%) P (%) S (%) Ni (%) Cr (%) Fe (%)

0.15584 0.17774 0.45330 0.17975 0.06918 0.1324 0.01567 98.8413

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moved along the weld area, a constant amount ofheat flux based on heat input was applied with time(depending on the welding speed) [11, 19]. The mov-ing heat load was applied on the area bounded by theweld lines, as shown in Fig. 1, and except for the weldzone other areas of the plates exposed to the atmo-sphere are assumed to be also subjected to heat lossby convection. In this analysis the convection losswas taken as 15 W/m2 K [7].

The following assumptions were made in the finiteelement analyses:

(a) density is not affected due to thermal expansion;(b) linear Newtonian convective cooling was

assumed;(c) convective cooling was assumed on all the sur-

faces except the weld zone;

(d) the heat source was assumed to have a Gaussiandistribution of heat flux;

(e) arc energy transfer efficiency (h ¼ 0.90) [20] wastaken to account for other losses.

The governing differential equation for heatconduction in a solid without heat generation isgiven by

@

@xk@T

@x

� �þ @

@yk@T

@y

� �þ @

@zk@T

@z

� �¼ rc

@T

@tð2Þ

3.1.1 Boundary condition

The finite-element-based weld thermal modellingcarried out in the present investigation is governed

Fig. 1 Schematic diagram of a square butt joint (numerical numbers representing corner points andpositions of tacks)

Fig. 2 Modelling and meshing of double-pass butt joints considering bead width and top and bottomreinforcements

394 M M Mahapatra, G L Datta, B Pradhan, and N R Mandal

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by equation (2) and is subjected to three boundaryconditions.

(a) First boundary condition. A specified initialtemperature for the welding that covers all the finiteelements of the model

T ¼ T1 for t ¼ 0 ð3Þwhere T1 is the ambient temperature. The energybalance at the work surface leads to the second andthird boundary conditions.

Let S1 represent the weld zone where the heat fluxof the welding arc calculated according to equation(1) is applied. The remaining part of the plate surfacerepresented by S2 is exposed to atmosphere whereheat loss also takes place due to convection. Theheat flux supplied to the model and convection lossto the atmosphere are indicated in the second andthird boundary conditions respectively.

(b) Second boundary condition. Heat flux due tothe welding arc is applied over the surface of theweld zone

qn¼ � qsup

or

�k@T

@n¼ �qsup on the surface S1 for t > 1 ð4Þ

(c) Third boundary condition. Convection loss isapplied to the surface except for the weld zone inthe finite element model

�k@T

@n¼ �hfðT1�T Þ on surface S2 for t > 0 ð5Þ

To avoid the sharp change in the value of specificheat, enthalpy was used as the material property.This was done by defining enthalpy of the materialas a function of temperature [11, 19]. To simulate fil-ler material deposition for each welding pass the ele-ment birth and death technique was used. For this,the elements were deactivated and activated as theheat source moved along the weld line during eachpass of welding. Before applying heat flux to the ele-ments of the first pass all the elements of the secondpass were deactivated.

3.2 Structural model

The stress–strain relationship can be represented as

fsg ¼ ½D�f«eg ð6Þwhere

f«eg ¼ f«g�f«tg ð7Þand

f«tg ¼ DT ½ax ay az 0 0 0�T ð8Þ

whereDT ¼ Tn�T1 and Tn is the instant temperature at

the point of interest.Considering the plastic strains, equation (7) can be

written as

f«eg ¼ f«g�f«tg�f«pg ð9ÞTransient thermal and non-linear structural analy-

sis was done for predicting angular distortion. Foreach pass of welding the temperature history fromthe thermal analysis in every load step was used asthe thermal loading in the structural analysis. Thestructural analysis involves large displacements(strain) and a rate-independent thermoelastoplasticmaterial model with temperature-dependent ma-terial properties incorporated into the modelling.Kinematic work hardening together with the vonMises yield criterion and associative flow rules [4, 5,21] were assumed in the analysis. In the structuralanalysis boundary conditions that prevented rigidbody motions were imposed into the modelling. Inthe present work eight-noded brick elements wereused for the thermal analysis and similar eight-nodedelements were used in the structural analysis.The eight-noded brick elements were chosen forrequired good compatibility in thermomechanicalanalysis. The solution was obtained using the ANSYSpackage [22].

4 EXPERIMENTAL DETAILS

Submerged arc welding of test samples was carried outwith various combinations of welding speed, current,and voltage. Chromel–alumel thermocouples wereused on the bottom side of the plates to record thetemperature history during welding. A typical weldingset-up is shown in Fig. 3. A double-pass submerged

Fig. 3 Submerged arc welding in progress

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arc-welded square butt joint marked with grid pointsto measure angular distortions with first and secondpass welds is shown in Fig. 4. The length and widthof each plate was 220 and 100mm respectively.

All plates were welded in the square butt conditionwithout any edge preparation in two passes. The typeof filler wire used was AWS SFA 5.17 ELB of 4mm dia-meter. The type of flux used conformed to AWSA5.17-80/SFA 5.17, having a basicity index of 1.6.The plates were cleaned and tack-welded with suita-ble root gaps before welding. The welding and otherparameters of a sample of two testpieces are givenin Table 3. Each joint was tack-welded at two points.Each tack point was situated 5mm away (at points 9and 10 in Fig. 1) from the edge (end) of the plate.The grid points were marked on the tack-weldedplates. An experimental set-up on a machine bedfor measuring the displacements of the grid pointsalong the vertical axis of the butt joints is shown inFig. 5. To facilitate clearance of the plate and the

weld from the machine bed surface, the joints wereplaced on two marking blocks, as shown in Fig. 5(a).The linear variable differential transducer (LVDT)probe was attached with the vertical axis of themachine shown in Fig. 5(b). After the welding, thedisplacements of each grid point with respect to itspre-welding readings were noted. The vertical displa-cements were compared with those obtained fromthe finite element analysis.

4.1 Achieving top and bottom reinforcements

In an automatic welding process like SAW, setting thewelding parameters is the most important step inachieving the desired results [23]. Several experi-ments were conducted to observe the effects of pro-cess parameters on weld metal deposition andreinforcements. Instead of using a costly ceramicbacking strip, a specially designed reusable alu-minium backing strip was used in the experimentsfor supporting flux and molten metal at the bottomside of the joints. For jobs 1 and 2 (12mm thickplates) the root gap maintained was 3.5mm. Theinput current and voltage were varied for jobs 1 and2 for each pass during experiments for obtaining dif-ferent bottom and top reinforcements. In the firstwelding pass the bottom reinforcement and fillingup of the root gap were achieved. In all the casesthe electrode feed rate was more than 2.75m/min.Good bottom reinforcements were achieved by vary-ing current and voltage (Table 3) in the ranges of

Fig. 4 A double-pass submerged arc-welded square buttjoint

Table 3 Specification of welding

Jobnumber

Platethickness(mm)

Current in firstpass (A)

Current in secondpass (A)

Voltage infirst pass (V)

Voltage in secondpass (V)

Speed in firstpass (mm/s)

Speed in secondpass (mm/s)

1 12 580 365 31 30 5.64 8.52 12 550 340 28 29 5.64 8.5

Fig. 5 Experimental set-up for measuring the grid point displacements of a submerged arc-weldedsquare butt joint: (a) schematic diagram of the measuring set-up and (b) measuring the dis-placements of grid points

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550 to 590A and 28 to 30V respectively. In the secondwelding pass different top side weld reinforcementswere intended for jobs 1 and 2. The second pass ofwelding started after 180 s of cooling of the first passweld. The time was sufficient for removing the fluxand moving the trolley to the initial starting positionof welding. A high welding speed and low currentwere used in the second pass to provide good topside bead reinforcement.

4.2 Measurement of zones of microstructures

For studying the microstructures of different zonesthe welded joints were sectioned and metallographicsamples were prepared [24] and examined under anoptical microscope. The measured zones and macro-graphs of jobs 1 and 2 are shown in Fig. 6. Bottomreinforcement, fusion zone (FZ 1), and heat affectedzone (HAZ) (HAZ 1) due to the first pass are shownin Fig. 6, as well as reinforcement, fusion zone (FZ 2),and HAZ (HAZ 2) due to the second pass.

5 MODELLING AND SOLUTIONS

Three-dimensional transient thermal analyses werecarried out for 12mm thick double-pass square butt

joints (jobs 1 and 2). The length and width of thesquare butt models were the same as those in theexperiments. As in the experiments for jobs 1 and 2,moving heat flux was applied at the arc traverse rateof 5.64mm/s for the first weld pass and the subse-quently cooling was applied for up to 220 s. For thesecond weld pass, moving heat flux was applied atthe arc traverse rate of 8.5mm/s and subsequentlycooling was applied for up to 520 s.

After obtaining the temperature distributions ofsquare butt welds and matching the trends with themicrostructure zones, transient thermal and non-linear structural analyses were carried out for predict-ing the angular distortions. The arc radius of each jobwas selected on a trial-and-error basis [11, 19]. Var-ious arc radii were tried in order to obtain the appro-priate temperature distributions that matched withthe experiments. The arc radii used in the analysisfor the first and second passes of jobs 1 and 2 weremeasured as 6.3, 6.2 and 6.1, 6.15mm respectively.The top and bottom reinforcement heights of jobs 1and 2 (Fig. 6) were measured as 2.18, 0.96 and 2.01,1.01mm respectively. The top and bottom reinforce-ment widths of jobs 1 and 2 were measured as 16.8,5.6 and 18.8, 9.4mm respectively. These readingswere incorporated into the modelling.

Fig. 6 Measured zones and macrographs of jobs 1 and 2: BW1, bead width of bottom reinforcement; BW2,bead width of top reinforcement; BH1, bead height of bottom reinforcement; BH2, bead height of topreinforcement; FZ 1, fusion zone due to the first pass; FZ 2, fusion zone due to the second pass: (a)zones of double-pass butt welds, (b) macrograph of job 1, and (c) macrograph of job 2

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6 RESULTS AND DISCUSSIONS

Modelling and meshing of a 12mm thick double-passbutt joint with top and bottom reinforcements areshown in Fig. 2. The weld zone was finely meshedand a gradual increase in mesh size was incorporatedaway from the weld line. Predicted top side tempera-ture distributions for the first pass at 150mm fromthe edge 5–6–7–8 (Fig. 1) along the Z direction (i.e.from start of welding) of job 1 are shown in Fig. 7.The figure shows the peak temperature rise to about2200 �C and the heating and cooling curves at theweld central line. Predicted top side temperature dis-tributions for the second pass at 150mm from theedge 5–6–7–8 (Fig. 1) along the Z direction (i.e. fromthe start of welding) (Fig. 2) of job 1 are shown inFig. 8. In Fig. 8 the initial temperature of the second

weld pass is predicted as the temperature to whichthe first pass weld had cooled down at 220 s. Thepeak temperature attained during the second pass isless in comparison to the first pass welding becauseof the high arc traverse speed used for the secondpass welding. It can also be observed that moregradual cooling took place in the second pass weld(Fig. 8) as the plates were already pre-heated due tothe first pass welding (Fig. 7). A comparative bottomside temperature distribution at 36mm away fromthe weld line obtained experimentally as well as byusing the theoretical model is shown in Fig. 9. Fairlygood agreement can be observed between the mea-sured and numerically obtained temperaturedistributions, indicating adequacy of the modeldeveloped. The predicted top side temperature distri-butions for the second pass of job 2 at 150mm fromedge 5–6–7–8 (Fig. 1) (i.e. from the start of welding)in the welding direction is shown in Fig. 10. Higher

Fig. 7 Predicted top surface nodal temperature history ofan element belonging to the first weld pass of job1 with cooling up to 220 s

Fig. 8 Predicted top surface nodal temperature history ofan element belonging to the secondweld pass of job 1

Fig. 9 Comparison of bottom side temperature distribu-tion 36mm away from the weld line of job 1

Fig. 10 Predicted topsurfacenodal temperaturehistoryof anelement belonging to the second weld pass of job 2

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peak temperature distributions were predicted for job1 due to a higher heat input condition.

This investigation is limited to the prediction ofangular distortion only. Here the displacement ofgrid points in the Y direction in each pass of weldingis considered as angular displacement and is com-pared with nodal displacement obtained from thefinite element model in the Y direction. The min-imum and maximum angular distortions obtained

from the finite element analysis and experimentsare shown in Table 4. The maximum angular deflec-tions in the Y direction, represented as SMX in mfor the first pass welding, are shown for job 1 inFig. 11. The maximum angular deflections in the Ydirection, represented as SMX in m after the secondpass welding, are shown for jobs 1 and 2 in Figs 12and 13 respectively. As expected, a greater amount

Table 4 Predicted and measured peak distortions

Job numberPlate thickness(mm)

Predicted minimumangular distortion(mm)

Measured minimumangular distortion(mm)

Predicted maximumangular distortion(mm)

Measured maximumangular distortion(mm)

1 12 0.98 0.7 1.32 1.22 12 0.91 0.7 1.21 1.3

Fig. 11 First pass angular distortion of job 1 expressed as adisplacement (SMX in m) in the Y direction

Fig. 12 Second pass angular distortion of job 1 expressedas a displacement (SMX in m) in the Y direction

Fig. 13 Second pass angular distortion of job 2 expressedas a displacement (SMX in m) in the Y direction

Fig. 14 Measured and predicted displacements of points(SMX in m) in the Y direction for job 1 along theedge parallel to the weld line

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of angular distortion is observed in the case of job 1because of more heat input. The comparisons ofdeformations obtained from modelling and experi-ments at the plate edge parallel to the weld line forjobs 1 and 2 are shown in Figs 14 and 15 respectively.Although the data for job 1 show a fairly close matchbetween the experimental and the modelled values,the same cannot be said for job 2 for which the rateof heat input is less. This difference may be due toexperimental error.

7 CONCLUSIONS

The following can be stated from the present experi-mental and modelling investigations on double-passsubmerged arc-welded butt welding.

1. Experiments conducted with proper weld passparameters and flux-filled aluminiumbacking stripled to adequate bottom reinforcement, root gapfill-up in the first pass, and required topreinforcement in the second pass. This ensuredthat theweld in the first pass acted like a constraintin the second pass, leading to less angular distor-tion.

2. A three-dimensional finite element model forpredicting the temperature distributions andangular distortions of double-pass submergedarc-welded square butt joints has been developed.

3. Experimental temperature distributions of thesquare butt joints away from the weld line closelymatched the values obtained from the finiteelement modelling.

4. Layer-wise application of heat flux, incorporationof joint geometry into the modeling, andconsideration of filler material deposition in the

analysis led to temperature distribution profilesthat closely matched the experimental values.

5. The angular distortions of the plates were mod-elled by three-dimensional finite element analy-sis and the results compared with theexperimentally measured values.

REFERENCES

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Fig. 15 Measured and predicted displacements of points(SMX in m) in the Y direction for job 2 along theedge parallel to the weld line

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APPENDIX

Notation

c specific heat (J/kg K)[D] stress–strain matrixhf convective heat transfer coefficient

(W/m2K)k thermal conductivity (W/mK)qconv heat loss from the work surfaces by

convection (W/m2)qn component of conduction heat flux

normal to the work surface (W/m2)qsup supplied heat flux from the welding arc

(W/m2)Q arc power (W)r radial distance in the welding arc (mm)S1 and S2 surfacest time (s)T temperature (�C)T1 temperature of the surroundings (�C)x, y, z coordinates

ax, ay, az thermal expansion coefficients in the x, yand z coordinates

{«} total strain vector{«e} elastic strain vector{«t} thermal strain vector{«p} plastic strain vectorh arc energy transfer efficiencyr density of the base metal (kg/m3){s} stress vector

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