JWJ > Volume 36(2); 2018 > Article
Narang, Mahapatra, Jha, Sridhar, and Biswas: Experimental and Numerical Study on Effect of Weld Reinforcement on Angular Distortion of SAW Square Butt Welded Plates

### Abstract

Weld reinforcements play vital role in controlling different formation of inherent strains in the heat-affected zone in arc welded joints. It is desirable to have adequate weld reinforcements on both the top and bottom sides of butt joint such that shrinkage force effects of both the reinforcements would be cancelling the angular distortion. In this study, it has been established by experiments that acceptable top and bottom weld reinforcements can be achieved and angular distortion can be controlled in square butt joints through submerged arc welding (SAW) done from both the sides with accurate control of welding process parameters. Such process does not require use of costly backing strip and instead a flux bed can be used for the purpose. Modeling of the process was also carried out using 3-D thermomechanical finite element (FE) analysis. Moving distributed heat source, weld reinforcements on top and bottom side, weld deposition in each welding pass and temperature dependent thermos-physical properties were used in the thermomechanical analysis for predicting the temperature profiles and angular distortion patterns The numerically predicted temperature distribution and distortion patterns indicated close agreement with the experimentally observed results.

### 1. Introduction

This present study dealt with single pass double side SAW square butt welding of 12 mm thick structural steel plates by using a flux bed. Use of costly backing strip was thus eliminated The SAW square butt joints were welded from both side sequentially. The idea was to have adequate weld reinforcement on both the sides of square butt joints such that the shrinkage forces on both the sides would cancel out each other resulting in minimized angular distortion. The phenomenon is shown in a simple schematic in Fig. 1 (a) and (b) which indicates cross sections of two square butt welds with different weld reinforcements. Fig. 1 (a) indicates greater top side shrinkage force (TSF) due to larger weld deposition at the top side leading to higher angular distortion. The topside shrinkage force (TSF) and bottom side shrinkage force (BSF) are nearly equal for the butt welded joint shown in Fig. 1 (b) resulting in lower angular distortion.
##### Fig. 1
(a) A butt joint with greater top side shrinkage force (TSF) due to greater weld deposition at the top side; (b) A butt joint with near equal top side shrinkage force (TSF) and bottom side shrinkage force (BSF)

### 2. Modeling methodology

The 3-D finite element analyses of SAW butts were carried out considering the following: (1) heat source with movement as per welding speed, (2) temperature dependent thermos-physical properties, (3) joint geometry such as upper bead and lower bead reinforcements in the modeling, and (4) incorporation of suitable material model for the elastic-plastic behavior of the structural steel weld and base metal. The transient thermal history was compared with the experimental data such as thermocouple readings and macrostructure zones’ boundary. Transient thermal and non-linear elasto-plastic thermos-mechanical analyses were performed to predict the temperature distribution in each welding pass and angular distortion of the SAW square butts. The developed methodology was also compared with the patterns of distortion obtained from experiment. The chemical composition of the base metal used for SAW is given in the earlier publication27).
The heat source used in this present work was considered as distributed heat flux. Varying welding parameters such as weld traverse rate, current and voltage were used in the experiment. The transient temperature data were measured on both top and bottom sides of the plate using thermocouples for each welding pass away from the weld center. The temperature profiles obtained from transient thermal analysis were compared with measured thermocouple readings. The heat flux used in FE model was considered as a Gaussian distributed heat flux as stated in equation (1):
##### (1)
q(r)=3Qeπr¯exp[3(r¯r)2]
where r̄ is characteristic radius of heat source where 95% of heat flux is concentrated27-29) and
Qe= ηVI = power of welding arc, W
η = welding arc efficiency
I&V= arc current in ‘amp’ and arc voltage ‘volt’ re- spectively.

### 3. Thermal modeling of SAW bead-on-plates

The 3-D FE analysis of the SAW bead-on-plates were carried out considering the following: (i) moving co-ordinate heat source, (ii) temperature dependent material properties, (iii) bead geometry such as reinforcement on both sides of joints. The moving heat source model was initially tested for SAW butt bead-on-plates for predicting the thermal histories and macrostructure boundary zones. For SAW bead-on-plates, the temperature distribution were fairly matched with the experimentally obtained thermal profile such as thermocouple readings and macrostructure zones’ boundary. Once the moving heat source model’s effectiveness was established for SAW bead-on-plates, the modelingl technique was further extended for thermomechanica analysis of square butt joints.
The top side temperatures of bead-on-plate welds, at distance from the weld, were measured by K-type thermocouples. Due to the flux cover over SAW weld bead and arc it is impossible to measure the SAW arc spread. Hence arc radius of each but-bead-on plate weld was estimated11,17,27) based on the electrode diameter and dimension of weld bead reinforcement. The schematic of butt-bead-on plate is shown in Fig. 2.
##### Fig. 2
The thermos-physical material properties26-29) used in the analysis to estimate temperature profiles and distortions are presented in Fig. 3. The constant material density used in the analysis was taken as 7850 kg/m3 26-29). The liquidus (Tliquidus) and Solidus (Tsolidus) temperatures of the base material were considered as 1500° and 1450° respectively11).
##### Fig. 3
Material properties used in finite element analysis
Natural Newtonian convective cooling was assumed for all the surfaces of the welded joint except the weld region. To account for other losses, arc efficiency (h = 0.90) was taken into consideration11,27). The thermal governing equation in an isotropic solid is given by:
##### (2)
x[KTx]+y[KTy]+z[Ktz]=ρcTt

### 3.1 Boundary conditions

Initial condition: A specified room temperature was considered for the entire surface of welding plate
##### (3)
T=Tafortime't'=0
where Ta is the surrounding temperature. As the weld area is completely covered by flux granules in SAW, no convectional heat transferee was considered for the weld region. For second and third boundary conditions, the energy balance criteria has been considered at the work surface.
First boundary condition: The specified heat flow acting over the surface
##### (4)
{q}T{n}=qsupontheweldzonesurfaceS1fortime't'>0
where, (qsup) is a specified heat flow supplied from an external welding arc over the instantaneous surface S1 (weld zone) of work.
2nd boundary condition: Heat loss over the surface other than welding region S2 has been specified by Natural Newton law of convection which is as below:
##### (5)
qn=qconor{q}T{n}=hf(TT)onS2fortimet>0
where S2 represents the surface of work other than weld zone exposed and (qcon) = convection heat loss.
Temperature dependent enthalpy was used as one of the material property for avoiding the sharp change in the value of specific heat. The chemical composition of structural steel used in the experiment is given in Table 1. The process parameters of SAW bead-on-plates are indicated in Table 2. The SAW bead-on-plates FE modelling geometries with different weld reinforcements together with the etched weld cross-sections are shown in
##### Table 1
Composition of the structural steel used in the experiments
C% Si% Mn% P% S% Ni% Cr% Fe%
0.160 0.177 0.453 0.179 0.069 0.132 0.015 98.841
##### Table 2
Process parameters of SAW bead-on-plates and weld dimensions
Sample No. Welding Current (A) Welding Voltage (V) Traverse Speed [WS] (mm/sec) Bead Height (mm) Bead width (mm)
1 450 29 5.56 3.07 17.25
2 475 32 6.94 2.83 16.88
3 525 36 8.33 2.16 17.53

### 3.2 Result of thermal analysis of SAW bead-on-plates

The result of thermal analysis for Job 1 as per parameters stated in Table 2 is presented in Fig. 5 (a). It is observed that the peak temperature exceeds 2550° at the middle of the weld as indicated in Fig. 5(a). The temperature during welding was measured using K-type thermocouples. The thermocouples were connected to AGILENT 34972A data logger device with 20 channels multiplexer. This data logging device was connected to a PC and the temperature at different locations were measured & recorded with time. Time dependent temperature data during welding at 22 mm from the weld center were noted by thermocouples. The experimentally observed and numerically obtained temperature profiles 22 mm away from the weld centre are plotted in Fig. 5(b) and close agreement is observed between the two with a variation of 6% only in case of peak temperatures.
##### Fig. 4
The etched weldment crosses sections of: (a) Sample 1; (b) Sample 2; (c) Sample 3 and model & meshing of butt-bead-on with weld reinforcement of (d) Sample 1; (e) Sample 2; (f) Sample 3
##### Fig. 5
(a) Peak temperatures distribution of Job 1; (b) Comparison of temperatures profiles 22mm away from the weld centre; (c) Peak temperatures distribution of Job 2 and (d) Peak temperatures distribution of Job 3
The result of thermal analysis for Job 2 is presented in Fig. 5 (c). It is observed that the peak temperature exceeds 2781° at the middle of the weld as indicated in Fig. 5 (c). The temperatures observed for Job 2 is higher than that of Job 1 due to higher current and voltage. Similar trend was also observed for Job 3 in Fig. 5 (d) where the peak temperature attains higher values than that of Job 1. The temperature distributions obtained from the model was analysis to predict the boundary of fusion zone in terms of weld width and depth of penetration. The numerically predicted and experimental measure zones based on the peak temperatures attained are given in Table 3.
##### Table 3
Numerical and experimental SAW bead-on-plate weld dimensions
Job No. Numerical peak temperature (° C) Measured bead width (mm) Estimated bead width (mm) % Error of bead width prediction Measured depth of penetration Estimated depth of penetration % Error of depth of penetration
1 2550 15.25 13.9 -8.85 4.25 3.95 -7.05
2 2781 16.88 13.68 -18.95 4.38 4.34 0.91
3 2714 17.53 14.44 -17.62 5.07 4.87 -3.94
It is also observed that the experimentally observed temperatures distributions fairly matched with and numerically predicted temperatures distributions. The moving distributed heat source based thermal modeling of SAW bead-on-plates proved to be successful in predicting the temperature distribution and macrostructure zones.

### 4.1 Experimental details

The experimental setup and the schematic of weld geometry cross-section are shown in Fig. 6 (a) and (b). The specification of the filler wire used was AWS EH-14, Grade C, 4 mm diameter. The SAW flux used was granular, basic and fluoride. The tack welding was applied at two points which are 5 mm away from the both edge (ends) of the plates. The welding was done from both sides of plates.
##### Fig. 6
(a) Experimental setup of SAW; (b) Schematic of a single pass double side SAW weldment indicating the bead widths and reinforcement heights (BH 1 & BH 2) of each weld pass
In order to avoid the occurrence of slag inclusion in the joint, it was required that in each pass of welding more than 50% of depth of penetration in terms of plate thickness (12 mm) is achieved. To accomplish the task a series of trial experiment runs were made and penetration achieved was measured as stated in Table 4. It is observed from Table 4 and Fig. 7 that with 32 V, 600 A and welding speed of 5.5 mm/sec more that 50% depth of penetration in terms of plate thickness could be achieved. Since the depth of fusion is less than 70% of the joint thickness, no extra arrangement such as backing strip was required to support the molten weld metal. Instead a flux bed was quite suitable for the purpose. This minimized the production cost. After each pass of welding the root was cleaned of slag with a wire brush and welding continued for the second side pass. Based on the trial experiments, the parameters of square butts were decided and presented in Table 5.
##### Table 4
Trial experimental runs for achieving depth of penetration more than 50% of plate thickness
Trial No. Voltage (volt) Current (amp) WS (mm/s) Depth of penetration in % of thickness
1 32 525 6.9 41.708
2 32 550 5.55 46.167
3 32 600 5.55 56.95
##### Fig. 7
Effect of process parameters on (DP) depth of penetration: (a) & (b) Depth of penetration of Trail 1 &2 less than 50% of plate thickness leading to entrapped slag in the joint; (c) Depth of penetration more than 50% of plate thickness for Trial 3 (Table 4)
The process is schematically shown in Fig. 8. The dimensions of each plate were 300 mm × 100 mm × 12 mm. The plate surface was marked with grid points along the plate edges as shown in Fig. 8 to measure the angular distortion. In the finite element analysis, the displacements of these points are regarded as the nodal displacement. The two welding heat sources used in both the weld passes on both sides are also shown in Fig. 8. Schematic of a single pass double side SAW butt weld in shown in Fig. 6 (b). The single pass double side SAW square butt joint process parameters are stated in the Table 5. To measure the distortion, the specified locations were marked on the tack-welded plates. A linear variable transformer (LVDT) was used for measuring grid point displacement. After completion of welding, the displacements of predefined marked locations were noted. The experimental distortions were calculated by subtracting the pre-welding readings from after welding readings. The vertical displacements of the marked locations of the joints were compared with the FE analysis results
##### Table 5
Single pass double side SAW square butt joint process parameters
Sample No Current (amp) Voltage (volt) WS (mm/s) BW1 (mm) BW2 (mm) H1 (mm) H2 (mm) Total area (mm) Maximum distortion(mm)
1 600 30 7.22 17.5 18.69 1.6 2.38 214.12 0.532
2 615 32 7.78 20.11 20.08 1.72 2.43 213.95 0.467
3 625 34 7.78 22.21 21.5 1.73 2.65 207.32 0.621
##### Fig. 8
Schematic of single pass double side SAW square butt joint with two moving heat sources and grid points for measuring distortion

### 4.2 Thermal analysis of single pass double side SAW square butt joints

The procedure of moving Gaussian distributed heat source adopted for the FE thermal analysis of SAW bead-on-plates proved it’s adequacy in predicting the temperature distribution and boundaries of fusion zone. The model was further extended for predicting the temperature distributions in case of single pass double side SAW square butt joints. The idea was to achieve adequate weld reinforcement on both sides of the plates to minimize angular distortion and to avoid the use of costly backing strip. Since in SAW the spread of arc is not visible, it has to be estimated. In the present investigation the arc radii were estimated based on the weld reinforcement dimension and subsequently used in the modeling for the prediction of transient temperature distribution11,27). The resulting temperature profiles were further compared with the experimentally obtained profiles. In the FE analysis the moving heat flux was applied as given in equation (1). Here the arc energy transfer efficiency (\$) of 90% was used11,27). The boundary conditions used for the SAW square butt thermal modeling is as described in equations 2-5. The polished and etched weldment cross sections of Jobs 1-3 are presented in Fig. 9. The meshing and model of the jobs are presented in Fig. 10. The weld zone is finely meshed with gradual coarsening of meshing away from the weld zone. In the FE model, the dimensions of welds (weld width and reinforcement) were incorporated as indicated for Jobs 1-2 in Fig. 10. Eight node brick element was used in the thermal model. For the structural model eight node brick element was also used for good compatibility.
##### Fig. 9
Macrostructure of SAW welds: (a) Weld cross section of Job 1 (Table 5); (b) Weld cross section of Job 2 (Table 5); (c) Weld cross section of Job 3 (Table 5)
##### Fig. 10
Modeling the SAW square butt joints according to the weld reinforcement height and width: (a) Model and meshing of Sample 1; (b) Model and meshing of Sample 2
The peak temperatures distribution for the first pass (first side weld) of Sample 1 (Table 5) is shown in Fig. 11 (a). It is observed that the temperatures rises up to 2000° very rapidly and then cools down till 450 seconds up to nearly 250° and then the second pass welding starts. The fist pass weld is heated up to 700° from 250° soon after the second side welding starts as indicated in Figs. 11 (a). This is the post welding heat treatment of first pass weld by the second pass welding. The peak temperatures distribution of second side welding is shown in Fig. 11 (b) for Sample 1. The dashed line portion of the temperature curve in Fig. 11 (b) indicates preheating effect of first pass welding. Due to the preheating effect of first pass welding, the second pass weld experience a temperature rise and welding starts at around 250° instead of room temperatures. Comparison of experimental and numerically predicted temperature profile at 22 mm away perpendicular to the centre of weld of Sample 1 is shown in Fig. 11 (c). The temperature during welding was measured using K-type thermocouples. The thermocouples were connected to AGILENT 34972A data logger device with 20 channels multiplexer. It can be observed that the experimental temperature profiles agree fairly well with the predicted ones with a variation of 11% only in case of peak temperatures.
##### Fig. 11
(a) Peak temperature distribution along the first pass weld line of Job 1 due to first and second side welding, (b) Peak temperature distribution of Job 1 along the second side weld line due to second weld pass (the preheating effect of first side welding shown in dashed line) and (c) Comparison of experimental and numerically predicted temperature profile 22 mm from weld centre of Job 1
The peak temperatures distribution for the first pass (first side weld) of Sample 2 is shown in Fig. 12 (a). It is observed that the temperatures shoots up above 2000° rapidly and then cools down till 450 seconds up to nearly 300°. The fist pass weld is heated up to 700° soon after the second side welding starts as indicated in Fig. 12 (a). The peak temperatures distribution of second side welding is shown in Fig. 12 (b) for Sample 2. The dashed line portion of the temperatures curve in Fig. 12 (a) indicates preheating effect of first pass welding. Due to the preheating effect of first pass welding, the second pass weld experience a temperature rise and welding starts at around 300° instead of room temperatures. Similar trend of first pass and second pass welding for Sample 3 are shown in Fig. 12 (c) & (d) respectively. For Sample 1-3 (Table 5) the experimental and numerical temperature profiles compared fairly well.
##### Fig. 12
(a) Peak temperature distribution along the first pass weld line of Job 2 due to first and second side welding, (b) Peak temperature distribution of Job 2 along the second side weld line due to second weld pass (the preheating effect of first side welding shown in dashed line), (c) Peak temperature distribution along the first pass weld line of Job 3 due to first and second side welding and (d) Peak temperature distribution of Job 3 along the second side weld line due to second weld pass (the preheating effect of first side welding shown in dashed line)

### 4.3 Structural analysis of single pass double side SAW square butt joints

The welding time was divided into several load steps in the thermal analysis and the resulting thermal condition was considered as the load for the sequential non-linear elasto-plastic structural analysis. Further, in the structural analysis, kinematic work hardening was considered for the weld and base metal4,17,27,28). Von Mises yield criterion and associative flow rule was also incorporated in the structural analysis4,17,27,28). 3-D 8-noded brick element was used in the thermal model. For the structural model eight node brick element was also used for good compatibility. The solution was obtained using FE package ANSYS30). The nonlinear thermo-mechanical analysis involved large displacements. In the structural analysis boundary conditions which prevented rigid body motions were imposed.
The distortion of Sample 1 expresses as the nodal deflection in ‘Y’ direction is shown in Fig. 13 (a). The nodal deflection patterns from the distortion model were compared with the grid point displacements of square butt joints. The nodal deflection in Y direction for Sample 2 and Sample 3 are shown in Fig. 13 (b) and (c) respectively. The comparative grid point displacement and nodal point deflection in Y direction of Sample 1 is presented in Fig. 14 (a). The comparative grid point displacements and nodal point deflections in Y direction of Job 2 and 3 are given in Fig. 14 (b) & (c) respectively. It can be observed that there is close agreement between the measured and predicted value of distortion. Maximum distortion of 0.558 mm was predicted for Sample 3 while minimum angular distortion of 0.430 mm was predicted for Job 2. It is observed from the process parameter Table 5 that Sample 1 is having least values of welding current, voltage and traverse speed. The heat source in Sample 1 travels at less speed in comparison to Sample 2 and 3. This might be the reason for higher angular distortion of Sample 1 when compared to that of Sample 2. Maximum angular distortion was exhibited in Sample 3 than Samples 1 & 2 due to higher welding current and voltage. Compared to single side square butt joints13,27) the single pass double side square butt joints exhibited lower angular distortion due to adequate weld reinforcement on both the side of the joint.
##### Fig. 13
(a) Distortion of Sample 1 expressed as displacement (SMX in m) in Y direction, (b) distortion of Sample 2 (SMX in m) in Y direction and (c) distortion of Sample 3 (SMX in m) in Y direction
##### Fig. 14
(a) Comparison between experimentally observed grid point displacement (distortion) and nodal displacement (SMX in m in Y direction) for Sample 1, (b) Comparison of experimentally observed grid point displacement (distortion) and nodal displacement (SMX in m in Y direction) for Sample 2 and (c) Comparison of experimentally observed grid point displacement (distortion) and nodal displacement (SMX in m in Y direction) for Sample 3

### 5. Conclusions

From the thermomechanical analysis of single pass double side SAW square butt joints following conclusions can be drawn:
• A 3-D transient elasto-plastic thermo-mechanical FE model was successfully developed to study the thermal history and distortion of single pass double side SAW square butt joints by considering top and bottom reinforcement. The developed modeled matched fairly well with that of experimental results with an error of 6-11% only.
• To minimize the angular distortion, single pass double side SAW square butts can be both side welded with a possibility of slag inclusion in the joint. However, it is observed that with 32 V, 600 A and welding travel speed of 5.5 mm/s more than 6 mm of depth of penetration can be achieved which eliminates the possibility of slag inclusion.
• This model successfully predicted the post welding heating effect of second pass weld on the first side weld. It also indicted the preheating of second pass weld due to effect of first pass weld.
• By this technique the use of costly backing bar/strip can be eliminated for thick section.

### References

1. Kamala V, Goldak J.A. Error due to two dimensional approximation in heat transfer analysis of welds. Welding Journal. 72 (9) (1993), 440s–446s

2. Tso-Liang T, Chin-Ping F, Peng-Hsiang C, Wei-Chun Y. Analysis of residual stresses and distortions in T-joint fillet welds. International Journal of Pressure Vessels and Piping. 78 (2001), 523–538 https://doi.org/10.1016/S0308-0161(01)00074-6

3. Radaj D. Heat effects of welding. Springer-Verlag GmbH. (1992)

4. Michaleris P, DeBiccari A. Prediction of welding distortion. Welding Journal. 76 (4) (1997), 172s–80s

5. Muransky O, Hamelin C, Smith M, Bendeich P, Edwards L. The effect of plasticity theory on predicted residual stress fields in numerical weld analyses. Computational Materials Science. 54 (2012), 125–134 https://doi.org/10.1016/j.commatsci.2011.10.026

6. Lindgren L.E. Finite element modeling and simulation of welding Part 1:increased complexity. Journal of Thermal Stresses. 24 (2001), 141–92 https://doi.org/10.1080/01495730150500442

7. Lindgren L.E. Finite element modeling and simulation of welding Part 2:improved material modeling. Journal of Thermal Stresses. 24 (2001), 195–231 https://doi.org/10.1080/014957301300006380

8. Alberg H, Daniel B. Comparison of an axisymmetric and a three-dimensional model for welding and stress relief heat treatment. In: In:Proceedings of the eighth international conference on numerical methods in industrial forming processes-NUMIFOFM 2004; Columbus, OH, USA. 13-17 June 2004.

9. Park J.U, Lee H.W, Bang H.S. Effects of mechanical constraints on angular distortion of welding joints. Science and Technology of Welding and Joining. 7 (4) (2002), 232–239 https://doi.org/10.1179/136217102225004266

10. Schenk T, Richardson I, Kraska M, Ohnimus S. A study on the influence of clamping on welding distortion. Computational Materials Science. 45 (4) (2009), 999–1005 https://doi.org/10.1016/j.commatsci.2009.01.004

11. Mahapatra M.M, Datta G.L, Pradhan B, Mandal N.R. Modeling the effects of constraints and SAW process parameters on angular distortions in one-sided fillet welds, Proceedings of the Institution of Mechanical Engineers, Part B. Journal of Engineering Manufacture. 221 (2007), 397–407 https://doi.org/10.1243/09544054JEM617

12. Camilleri D, McPherson T.G. Procedural tacking fabrication influences on welding distortion, Proceedings of the International Conference on Advances in Welding Science & Technology for Construction, Energy & Transportation. In : Kocak M, editor. (2010), p. 403–408

13. Mahapatra M.M, Datta G.L, Pradhan B, Mandal N.R. Three-dimensional finite element analysis to predict the effects of SAW process parameters on temperatures and angular distortions in single pass butt joints with top and bottom reinforcements. International Journal of Pressure Vessels and Piping. 83 (2006), 721–729 https://doi.org/10.1016/j.ijpvp.2006.07.011

14. Kannengiesser T, Lausch T, Kromm A. Effects of heat control on the stress build-up during high-strength steel welding under defined restraint conditions. Welding in the World. 55 (2011), 07–08 58-65

15. Bhatti A.A, Barsoum Z. Development of efficient three-dimensional welding simulation approach for residual stress estimation in different welded joints. Journal of Strain Analysis for Engineering Design. 47 (2012), 539–552 https://doi.org/10.1177/0309324712463866

16. Tsirkas S.A, Papanikos P, Pericleous K, Strusevich N, Boitout F, Bergheau J.M. Evaluation of distortions of laser welded shipbuilding parts using a local-global approach. Science and Technology of Welding and Joining. 8 (2) (2003), 79–88 https://doi.org/10.1179/136217103225010899

17. Mahapatra M.M, Datta G.L, Pradhan B. Three-dimensional finite element analysis to predict the effects of SMAW process parameters on temperature distributions and weldment zones in butt and one-sided fillet welds, Proceedings of the Institution of Mechanical Engineers, Part B. Journal of Engineering Manufacture. 220 (2006), 837–845 https://doi.org/10.1243/09544054JEM371

18. Fanous F.Z.I, Maher Y.A, Wifi S.A. 3-D Finite element modeling of the welding process using element birth and element movement techniques. Transactions ASME. Journal of Pressure Vessel Technology. 125 (2003), 144–150 https://doi.org/10.1115/1.1564070

19. Wang J, Ma N, Murakawa H, Teng B, Yuan S. Prediction and measurement of welding distortion of a spherical structure assembled from multi thin plates. Materials and Design. 32 (2011), 4728–4737 https://doi.org/10.1016/j.matdes.2011.06.047

20. Deng D, Murakawa H. Prediction of welding distortion and residual stress in a thin plate butt-welded joint. Computational Materials Science. 43 (2) (2008), 353–365 https://doi.org/10.1016/j.commatsci.2007.12.006

21. Deng D, Murakawa H, Liang W. Numerical simulation of welding distortion in large structures. Computer Methods in Applied Mechanics and Engineering. 196 (45) (2007), 4613–4627 https://doi.org/10.1016/j.cma.2007.05.023

22. Deng D, Murakawa H, Liang W. Prediction of welding distortion in a curved plate structure by means of elastic finite element method. Journal of Materials Processing Technology. 203 (2008) 252–266 https://doi.org/10.1016/j.jmatprotec.2007.10.009

23. Ueda Y, Murakawa H, Ma N. Chapter 2-Intro- duction to measurement and prediction of residual stresses with the help of inherent strains, Welding Deformation and Residual Stress Prevention. Boston. Butterworth Heinemann. (2002), 35–53

24. Murakawa H, Deng D, Ma N, Wang J. Applications of inherent strain and interface element to simulation of welding deformation in thin plate structures. Computational Materials Science. 51 (1) (2012), 43–52 https://doi.org/10.1016/j.commatsci.2011.06.040

25. Wang J, Shibahara M, Zhang X, Murakawa H. Investigation on twisting distortion of thin plate stiffened structure under welding. Journal of Materials Processing Technology. 212 (8) (2012), 1705–1715 https://doi.org/10.1016/j.jmatprotec.2012.03.015

26. Brown S, Song H. Implication of three-dimensional numerical simulations of welding large structures. Welding Journal. 71 (2) (1992), 55s–62s

27. Mahapatra M.M, Datta G.L, Pradhan B, Mandal N.R. Modelling of angular distortion of double-pass butt- welded plate Proceedings of the Institution of Mechanical Engineers, Part B. Journal Engineering Manufacture. 222 (2008), 391–401 https://doi.org/10.1243/09544054JEM995

28. Biswas P, Mandal N.R. Thermomechanical finite element analysis and experimental investigation of single-pass single-sided submerged arc welding of C-Mn steel plates, Proceedings of the Institution of Mechanical Engineers, Part B. Journal Engineering Manufacture. 224 (B4) (2010), 627–639 https://doi.org/10.1243/09544054JEM1624

29. Friedman E. Thermomechanical analysis of welding process using finite element method. Transaction of ASME Journal of Pressure Vessel Technology. 97 (3) (1975), 206–213 https://doi.org/10.1115/1.3454296

30. Theory reference, ANSYS Inc, Southpointe, Canonsburg, Pennsylvania. (2009)

TOOLS
Full text via DOI
CrossRef TDM
E-Mail
Print
Share:
METRICS
 0 Crossref
 664 View