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Journal of Materials Processing Technology 222 (2015) 344–355

Contents lists available at ScienceDirect

Journal of Materials Processing Technology

jo ur nal home p ag e: www.elsev ier .com/ locate / jmatprotec

ffect of material properties and mechanical tensioning load onesidual stress formation in GTA 304-A36 dissimilar weld

. Eisazadeha, A. Achuthana, J.A. Goldakb, D.K. Aiduna,∗

Department of Mechanical and Aeronautical Engineering, Clarkson University, 8 Clarkson Avenue, Potsdam, NY 13699, United StatesMechanical and Aerospace Engineering, Carleton University, Ottawa, ON K1S 5B6, Canada

r t i c l e i n f o

rticle history:eceived 29 September 2014eceived in revised form 13 March 2015ccepted 14 March 2015vailable online 24 March 2015

eywords:esidual stress

a b s t r a c t

Using a finite element analysis (FEA) model, the residual stress (RS) formation in an autogenous GTAdissimilar weld between austenitic stainless steel (304) and low carbon steel (A36) are analyzed. Theeffect of material properties on RS formation was determined by first considering a similar weld of 304plates, and then changing only a selected mechanical property of the 304 plate on one side of weld to thatcorresponding to an A36 plate. Enforcing one set of mechanical property to be different at a time helpedto isolate the role of these individual properties on the RS formation in the dissimilar weld. The effect ofmechanical tensioning on dissimilar welds is then investigated. Results show that the longitudinal RS in

imilar weldissimilar weldechanical tensioning

esidual stress reduction

both the similar and dissimilar welds can be reduced in the weld zone (WZ) by an amount equal to thestress corresponding to the applied mechanical tensioning load, as the tensioning load is removed aftercooling. The mechanism of RS formation in dissimilar weld, and its mitigation by mechanical tensioningare determined by comparing the longitudinal stress evolution on a cross-section of the dissimilar weldplates under the mechanically tensioned and free conditions.

. Introduction

Welding is a reliable and efficient metal joining process widelysed in the infrastructure and heavy equipment industry, such as

n the construction of steel bridges, shipbuilding, and installationf large pipelines. High joint efficiency, simple setup, and low fab-ication costs are the advantages of this joining process. However,ue to the localized heating, and subsequent cooling during weld-

ng, highly non-uniform temperature distribution occurs across theeld and the base metal (BM), resulting in the formation of sig-ificant residual stress (RS) in the weldment. The RS formation inelds, being a thermo-mechanical phenomenon that depends onany factors, makes its quantitative prediction quite challenging.

he principal factors that determine the RS formation in a weldedtructure weldment are shown in a Fishbone diagram in Fig. 1.nderstanding the distribution of RS induced by welding is crit-

cal in many industrial applications to determine the crack growthehavior and predict failure.

Several studies on the effect of RS on the failure of dissimilareld joint have been reported. For instance, Suzuki et al. (2012)

eported significant stress corrosion cracking (SCC) as a result of

∗ Corresponding author. Tel.: +1 315 268 6518.E-mail addresses: eisazah@clarkson.edu, heisazadeh@gmail.com (H. Eisazadeh),

aidun@clarkson.edu (D.K. Aidun).

ttp://dx.doi.org/10.1016/j.jmatprotec.2015.03.021924-0136/© 2015 Elsevier B.V. All rights reserved.

© 2015 Elsevier B.V. All rights reserved.

RS in the dissimilar welds between ferritic steels and austeniticstainless steels, which is widely used in the oil and gas industry. Ingeneral, the experimental determination of RS in dissimilar weldsis quite challenging when compared to the similar weld, due tothe differences in the BM properties, especially thermal expansioncoefficient and yield strength, producing a relatively complex dis-tribution of RS. Lately, a number of studies have used numericalmodels based on finite element analysis (FEA) to predict RS in dis-similar welds. For example, Ranjbarnodeh et al. (2011) determinedlongitudinal RS in butt joints of dissimilar steels and comparedthem with the stresses in similar steel joints. Similarly, Katsareasand Youtsos (2005) developed a 2-D FEA model relying on a simpleimplementation of the material property mismatch for RS predic-tion in dissimilar metal pipe welds. Lee and Chang (2007) studiedthe effect of yield and tensile strengths on RS by employing dif-ferent carbon steels under both similar and dissimilar butt weldconditions. Anawa and Olabi (2008) used FEA to determine theoptimized process parameters and develop statistical models forthe welding of stainless steel and low carbon steel using a CO2continuous-wave laser. Deng et al. (2009) determined the RS ina dissimilar metal pipe joint considering cladding, buttering, andpost-weld heat treatment (PWHT). Similarly, Lee and Chang (2011)

predicted the axial and hoop RS produced in high strength carbonsteel pipe weld using a FEA model by employing a sequentiallycoupled 3-D thermal and solid-state phase transformation duringwelding.

H. Eisazadeh et al. / Journal of Materials Processing Technology 222 (2015) 344–355 345

effect

tctiTitRiaibi(et

lrjsct(attrdtfistacw

a

Fig. 1. Fishbone diagram illustrates the

RS mitigation techniques have received increased attention overhe past two decades. Most of these techniques, however, are appli-able to similar weld. Zhang et al. (2005) have discussed the adapta-ion and optimization of some of these techniques, such as anneal-ng and peening processes, for its use on similar welded joints.here also exist specific welding methods such as in-process cool-ng and global mechanical tensioning (GMT), which are designedo work by influencing the thermal strain mechanism that drivesS formation in the weld. Altenkirch et al. (2008) experimentally

nvestigated the relationship between the final RS stress, distortion,nd the level of GMT applied in situ during welding. The mechan-cal tensioning technique, where a tensile stress is introducedefore welding and maintained throughout the welding and cool-

ng phases, was also found to be very effective by Richards et al.2008) in relieving the large longitudinal RS in similar welds. Went al. (2010) demonstrated that rolling the weld as an effective wayo improve the mechanical properties, and minimize distortion.

Studies with regard to mitigating RS in dissimilar welds areimited. Sedek et al. (2003) demonstrated that the thermal stresselieving techniques, such as the furnace annealing of weldedoints, were not only ineffective for mitigating RS in dissimilarteel welds, but actually increased RS as well. This is due to theonsiderable difference in the thermal expansion coefficient ofhe joined steels, producing large misfit stresses. Hurrell et al.2006) discussed a number of mechanical mitigation techniques in

review article. Broadly, they classified the mechanical mitigationechniques under three main categories: (a) surface mechanicalreatment to induce compressive skin stress, (b) mechanical stresselief through thickness, and (c) weld design optimization to pro-uce low/favorable RS levels and minimize distortion. Among allhe techniques in these three categories, Song et al. (2010) identi-ed weld overlay technique as the most effective for protecting atructural dissimilar metal weld. Kim et al. (2009) found that bothhe longitudinal and the transverse stress components decreaseds the number of layers increased. The effectiveness of mechani-

al tensioning technique has not been studied on a dissimilar weldhere the RS is not the same on the two base plates.

In summary, it can be stated that RS in dissimilar weld issymmetric and its highest magnitude occurs in the plate which

of various parameters on RS formation.

possesses greater yield strength and thermal expansion. In case offerritic steel and austenitic steel dissimilar weld, the latter under-goes higher RS. The belief behind this is that materials cannot holdRS beyond their yield point. Once stress magnitude at any locationof weld plate reaches to its corresponding yield point, that par-ticular location starts deforming plastically. Since austenitic steelhas greater yield point, the magnitude of RS in this metal will belarger. Thermal expansion of austenitic steel has a big role to playin this development, as well. Since this property is again greater inaustenitic stainless steel, area undergoing tensile RS will be larger.Overall, even though in the past decades, significant amount ofprogress on RS modeling was made, the mechanism of various mit-igation techniques, mechanical tensioning in particular, in terms ofits dependence on material properties of the two plates, is not wellunderstood. In the present study, we first used a FEA model to trackthe evolution of stress field at various locations during the weld-ing and cooling periods, providing a better understanding of themechanism of RS formation in dissimilar welds, particularly theinfluence of the difference in various material properties. Then themechanism of the influence of tensioning load on RS reduction insimilar and dissimilar welds was determined.

2. Methodology

In the experiment, the GTAW process without filler metal wasconsidered for three butt-welded joints of type 304 stainless steeland A36 low carbon steel. The welding parameters used for the sim-ulations conducted in this work are shown in Table 1. The chemicalcompositions of 304 and A36 plates are provided in Table 2.

The temperature fields and the evolution of the RS were inves-tigated by means of a sequentially coupled thermo-mechanicalformulation available in the ABAQUS commercial package. Thedetails of the model are shown in Fig. 2. A weld plate of dimen-sions 72 mm × 50 mm × 5 mm was considered. Linear 8-node brick

elements, with relatively finer elements in the 5 mm region on bothsides of the weld path, were used for the model. Simulation of weld-ing was realized by introducing a moving volumetric heat sourceregion into the plate at the desired speed in the desired direction.

346 H. Eisazadeh et al. / Journal of Materials Processing Technology 222 (2015) 344–355

Table 1Welding parameters used in this study.

Welding voltage (V) Welding current (A) Welding speed (mm/s) Arc efficiency Shielding gas, CFM (cubic feet per min)

11.5 100 2.4 70% Argon, 29

Table 2Chemical compositions (wt%) of A36 and 304 steels.

Grade Carbon Manganese Silicon Phosphorus Sulfur Chromium Nickel Balance

A36 0.08–0.29 0.40–1.20 0.15–0.40 0.04 Max. 0.05 Max. None None Fe304 Max 0.08 Max 2 Max 1 Max 0.045 Max 0.03 18–20 8–10 Fe

Fig. 2. Details of the plate geometry, mesh, and the thermal and mechanical bound-ale

2

hb

w

Q

Q

wmpafetdepcaiD

2

ta

Table 3Temperature dependent convection coefficients for the steel (Attarha and Sattari-Far, 2011).

h [W/m2 K] T − To [K]

1.85 569.079 278

ry conditions. T1, T2 and T3 represent the location of the thermocouples. Theocation P and cross-section A–A′ were used to study the temperature and stressvolution.

.1. Transient thermal analysis

The heat from the moving welding arc is applied as a volumetriceat source with a double ellipsoidal distribution and is expressedy the following equations (Goldak et al., 1984).

The ellipsoidal heat source distribution ahead and behind of theelding torch (arc) is given, respectively, as:

(x′, y′, z′, t) = 6√

3ff QW

a1bc�√

�e−3x′2/a2

1 e−3y′2/b2e−3z′2/c2

(1)

(x′, y′, z′, t) = 6√

3frQW

a2bc�√

�e−3x′2/a2

2 e−3y′2/b2e−3z′2/c2

(2)

here x′, y′ and z′ are the local coordinates of the double ellipsoidodel, relative to the torch location and aligned with the welded

lates. The ff and fr are the fraction of heat deposited in the frontnd the rear parts, respectively. Note that ff + fr = 2.0. In this study,

f is assumed to be 1.4 and fr to be 0.6, in order to introduce thexpected steeper temperature gradient in the ellipsoid, ahead ofhe torch. Qw is the power of the welding heat source, which can beetermined from the welding current, the arc voltage and the arcfficiency. The arc efficiency � is taken as 70% for the GTA weldingrocess (Deng and Murakawa, 2006). The parameters a1, a2, b and, define the characteristics of the welding heat source, and can bedjusted to create a desired melted zone according to the weld-ng conditions. The moving heat source is implemented using theFLUX user subroutine tool in ABAQUS.

.2. Mechanical analysis

The strain induced by welding process can be divided into elas-ic, plastic, thermal and transformation strain. The elastic stressnd strain was modeled using the isotropic Hooke’s law with

18.5 55652.6 2778

temperature-dependent Young’s modulus and Poisson’s ratio. Thethermal strain was computed using the temperature-dependentcoefficient of thermal expansion. For the plastic strain, a rate-independent plastic constitutive equation is considered with thevon Mises yield criterion, temperature-dependent mechanicalproperties and linear isotropic hardening rule. Since, the effect ofphase transformation for low carbon steel on welding deformationis insignificant and the phase transformation does not occur in theaustenitic stainless steel, it was ignored in the computational model(Deng et al., 2003).

2.3. Boundary conditions

The initial temperature of the material prior to welding was setto 298 ◦K. Thermal boundary conditions consist of the application ofconvective and radiative heat transfer to all surfaces of the model.Convective heat transfer coefficients are applied to the sides, top,and bottom surfaces of the plate as a function of the metal surfacetemperature (Table 3). The metal surface in the WZ and the sur-rounding region, being at high temperature, is dominated by heatlosses due to radiation, while losses due to convection dominate inthe relatively lower temperature surfaces, away from the weld zone(Deng and Murakawa, 2006). Table 3 presents the temperature-dependent convection coefficients used in our study. A value of800 W/m2 K [4–5] was considered for surfaces which are in contactwith the welding table.

Displacements are restricted on selected nodes (shown as 1,2 and 3 in Fig. 2) in order to prevent rigid body motion withoutconstraining the deformation of the weldment, as shown in Fig. 2(Biswas et al., 2011).

2.4. Material properties

The distribution of the heat in the metal depends on the mate-rial properties: thermal conductivity (k), specific heat capacity (C),mass density (�), latent heat of melting and latent heat of solid-ification. The temperature dependencies of these properties forboth 304 and A36 are shown in Fig. 3a. The latent heat of meltingand solidification is assumed to vary linearly between the solidus

and the liquidus temperatures. For 304, the latent heat, the solidustemperature and the liquidus temperature are 261 kJ, 1673 ◦K and1727 ◦K, respectively. Likewise, for A36, it is 247 kJ, 1738 ◦K and1817 ◦K, respectively.

H. Eisazadeh et al. / Journal of Materials Processing Technology 222 (2015) 344–355 347

Fe

dts

3

3

p3wucta

rasws

Fig. 4. Comparison between the measured and simulated temperatures for the sim-ilar welding of 304.

Fig. 5. Temperature distribution at the instance when the welding torch reaches themiddle of the weld path in 304 similar weld. (For interpretation of the references tocolor in the text, the reader is referred to the web version of this article.)

ig. 3. Properties of A36 and 304: (a) thermal properties, and (b) mechanical prop-rties.

Mechanical properties of 304 and A36, which are temperatureependent, are shown in Fig. 3b. The elastic–plastic material consti-utive model is implemented in the form of temperature dependenttress–strain curves.

. Results and discussion

.1. Experimental validation of the model

The FEA model was validated by comparing the predicted tem-erature with experimental data obtained for the similar weld with04 steel. The temperature histories at three different locationsere measured using K-type thermocouples spot welded to thepper and lower surfaces of the weld plate (Fig. 2). The thermo-ouples T1, T2, and T3 were placed at 3, 5 and 8 mm away fromhe weld line (boundary), respectively. The temperature data wascquired at a rate of 10 readings per second.

For the computational model validation, the temperature histo-ies predicted by the computational model throughout the heating

nd cooling stages were compared to the experimentally mea-ured values (Fig. 4). Overall, the model predictions agree wellith the measurements. Near the peak, agreement was relatively

tronger for locations T2 and T3, when compared to location T1.

Fig. 6. Longitudinal stress at the instance when the welding torch reaches the mid-dle of the weld path in 304 similar weld.

This could be attributed to the difference in the proximity of theselocations to the fusion boundary (FB). Being the nearest to the FB,T1 experiences the largest temperature change and the associatedproperty changes, thereby being the most sensitive to the underly-ing assumptions and input properties.

3.2. RS formation

Figs. 5 and 6 present the typical temperature and longitudinalstress distribution, respectively, on the surface of the plate duringwelding at the instance when the welding torch (arc) reaches the

348 H. Eisazadeh et al. / Journal of Materials Processing Technology 222 (2015) 344–355

Fig. 7. Longitudinal stress in 304 similar weld after cooling (longitudinal RS).

Fc

msealpsitstpprpita

lst(vAtt

Fig. 9. Temperature distribution at the instance when the welding torch reachesthe middle of the weld path in A36-304 dissimilar weld.

Fig. 10. Longitudinal stress at the instance when the welding torch reaches themiddle of the weld path in dissimilar weld of A36 and 304.

plate due to the relatively larger yield strength and thermal expan-

ig. 8. Temperature and longitudinal stress evolution at location P (Fig. 2) in theenter of weld plate in 304 similar weld.

iddle of the whole weld path. The temperature and longitudinaltress distribution are symmetric with respect to the weld path, asxpected. The temperature in the vicinity of the torch is very highnd, therefore, the material volume in this region is in a state ofow yield strength and, endures low stress. A large amount of com-ressive stress is generated just ahead of the torch because of theevere temperature gradient in this region, which is in the heat-ng phase. Severe temperature gradient also means the strength ofhe material increased quickly to room temperature value within ahort distance away from the weld zone (WZ). In the region behindhe weld torch (green region in Fig. 5), which is in the coolinghase, the temperature gradient gradually decreases as the tem-erature around the WZ continues to decrease, and it approachedelatively more uniform temperature distribution. As the coolingrogresses, the fusion zone material behind the torch contracts,

nitiating formation of a tensile RS. As the cooling completes, theensile longitudinal RS grows to a larger region throughout the WZ,s shown in Fig. 7.

The temperature and the stress history at an arbitrarily chosenocation in the WZ and in the middle of the weld path (location Phown in Fig. 2) are plotted in Fig. 8. During the initial heat up prioro the instance at which the center of the welding torch passes P0–10 s period), the severe localized temperature gradients in theicinity of the weld line created compressive stress in this region.

s the welding torch passed P, and the material around P started

o cool down, the large compressive stress transformed into a largeensile RS, consistent with the stress distribution discussed above.

Fig. 11. Longitudinal RS in dissimilar weld of A36 and 304 after cooling.

Fig. 9 displays the temperature distribution in a dissimilar weldwhen the torch reached the center of the weld path. As expected,the temperature distribution was asymmetric due to the differencein thermo-mechanical properties of the two plates. Heat diffusionduring the heating and cooling phases was relatively larger in A36due to its higher thermal diffusivity when compared to the 304.A36, with higher thermal conductivity than 304, also experienceda relatively faster temperature rise than 304. The longitudinal stressdistributions at this instance and after complete cooling are shownin Figs. 9 and 10, respectively. The stresses are higher in the 304

sion coefficient of 304 when compared to A36 (Fig. 11).The distribution of the longitudinal RS along the line A–A′

(shown in Fig. 2) after cooling is shown in Fig. 12 for a 304 (similar

H. Eisazadeh et al. / Journal of Materials Proce

Fig. 12. Longitudinal RS after cooling in similar A36, similar 304, and their combi-nation for dissimilar weld, which is shown by blue color, (left half is A36 and righthalf is 304).

Fa

wAmryAtdTwmtss

atfal

value in 304 reaching its yield strength. The results corresponding

ig. 13. Normalized Longitudinal RS (RS/YS) in dissimilar A36-304 weld (left-A36nd right- 304).

eld), A36 (similar weld) and their combination (dissimilar weld).s discussed above, for the same heat input and BM dimensions, theagnitude of RS in the dissimilar weld around the WZ (±7 mm) is

elatively higher in 304 than that in A36 due to the fairly largerield strength and thermal expansion of 304 when compared to36. The RS profiles of A36 and 304 similar welds, when compared

o that of their dissimilar weld (A36-304 weld), show a significantifference in the longitudinal RS magnitude, as shown in Fig. 12.he maximum value the longitudinal RS in 304 of the dissimilareld is 50 MPa more than that of the 304 similar weld, while theaximum longitudinal RS in the A36 is significantly lower than

he A36 similar weld. The longitudinal RS normalized with its yieldtrength (YS) is presented in Fig. 13 for the dissimilar weld, whichhows that the RS in 304 reached its yield strength in the WZ.

The formation of longitudinal RS in similar welds of 304, A36,nd the dissimilar welds of 304-A36 plates, is studied by capturinghe evolution of longitudinal stress distribution along line A–A′, for a

ew selected times (Fig. 14). After 3 s from the start (location showns 3 s in Fig. 14b), the longitudinal stress is very low due to theow temperature gradient since the welding torch (arc) is relatively

ssing Technology 222 (2015) 344–355 349

far from line A–A′ (Fig. 14). As the welding torch reaches the loca-tion corresponding to the 8 s instance from the start, compressivelongitudinal stress began to evolve, but mostly in the region closeto WZ. As the welding torch traveled forward and reached A–A′,and the temperature increased, a W shaped stress profile with arelatively low tensile stress in the WZ, large compressive stressnext to the WZ, and large tensile stress away from the WZ, devel-oped for both the similar and dissimilar welds. Subsequently, asthe welding torch crossed the section A–A′ and continued to travelforward, the tensile stress in the WZ increased in magnitude. Thecompressive stress magnitude remained the same, but spread intothe outer tensile region. The evolution of stress during the cool-ing stage, where thermal mismatch and the associated stressesgradually changed the non-uniform plastic deformation, eventu-ally resulted in a drastically different stress profile(@2500 s). Theouter region (±20 mm) with substantial tensile stress completelychanged to large compressive stress while the tensile stress in theWZ increased significantly, reaching the yield strength in 304.

3.3. Sensitivity of RS to various material properties

A sensitivity study was performed to identify the impact of thedissimilarity in various material properties of 304 and A36, namelythermal conductivity, specific heat capacity, yield strength, andcoefficient of thermal expansion, on the formation of RS. This wasachieved by characterizing the RS formation in the WZ of a 304plate, with the particular property under consideration changed tothe corresponding A36 property, but only for material on one sideof the weld (referred to as A36 plate in this section, as shown inFig. 15). This approach, where only a single material property wasdissimilar across the weld at a time, permitted to study the impactof individual material properties qualitatively (not quantitativebecause of elasto-plastic constitutive model which is nonlinear andhence history dependent) without being influenced by the dissim-ilarity of other material properties. The results are summarized inFig. 16. In case 1, since material properties are exactly the same onboth sides of the weld, longitudinal RS is symmetric with respectto the weld path, as discussed earlier. In case 2, bell shaped RSdistribution shifted slightly toward the right due to the differencein thermal conductivity between the plates. A36 has a much higherthermal conductivity coefficient than that of 304 in the tempera-ture range below 1000 K (Fig. 3). Since higher thermal conductivityleads to higher thermal diffusivity for a given mass density andspecific heat capacity ( ̨ = k/�Cp), heat diffusion is faster in the A36.Therefore, the temperature distribution is relatively more uniformin this region when compared to the locations away from the weldpath where temperature gradient is relatively large. In effect, thetemperature across the plate becomes uniform more quickly inA36. Therefore, the lower temperature gradient in the A36 plateinduced less tensile RS which was limited to a smaller region nearthe weld path. Since the specific heat capacity was similar for theA36 and 304 in the temperature range below 1000 K, the obtainedstress field for case 3 was very similar to case 1. The yield strengthwas varied in case 4, resulting in the lowering of RS in A36, whichhas the lower yield strength. This can be attributed to the depen-dence of the RS on the plastic strains in the WZ, which essentiallylimits the maximum stress to the yield strength of the weld mate-rial. Compared to thermal conductivity, specific heat, and yieldstrength, the effect of thermal expansion coefficient showed thelargest sensitivity to the RS in the WZ (case 5). Since the thermalexpansion coefficient of A36 was small compared to the 304, the RSdropped to a remarkably low value in A36, and increased to a larger

to the original dissimilar weld between A36 and 304, were alsoprovided as case 6 in Fig. 16 for comparison purposes. Interest-ingly, cases 5 and 6 produced very similar longitudinal RS states,

350 H. Eisazadeh et al. / Journal of Materials Processing Technology 222 (2015) 344–355

Fig. 14. (a) Predicted longitudinal RS profiles across the mid-plane of the weld at differeA36 and right half 304), (b) locations of welding heat source at different time. (For interpweb version of this article.)

Fit

idaw

ig. 15. Schematic illustrating how the variation in thermal expansion coefficientn a dissimilar weld can introduce mutual tension/compression loading conditionshat influence the RS formation.

ndicating that the coefficient of thermal expansion plays the mostominant role in the evolution of RS in the dissimilar welds of A36nd 304. Following the calculation shown by Deng et al. (2009),ithin the temperature range from room temperature (298 K) to

nt times (red square for similar weld, blue circle for dissimilar weld with left halfretation of the references to color in this figure legend, the reader is referred to the

807 K, the average difference of thermal expansion coefficient A36and 304 is almost 11 × 10−6. With this given difference, it can beconcluded a larger thermal strain will be generated in the 304 sidewhen the weld cools from 807 K to room temperature. Sedek et al.(2003) also stated the variation in the thermal expansion coefficientbetween ferritic–pearlitic and austenitic stainless steels in dissim-ilar weld caused RS differences in these two weld plates. HighestRS occurred in the latter because of its greater thermal expansioncoefficient. They suggested these high stresses could be reducedconsiderably by using a transition material having a coefficient ofthermal expansion between those of the two base metals.

Fig. 15 illustrates that during weld thermal cycle, the 304 platewith large thermal expansion coefficient can produce tensioningload on A36 plate, while enduring compression load by itself.

3.4. Effect of tensioning load on RS

Price et al. (2007) and Richards et al. (2008) have shown that the

application of global, or far field, mechanical tensioning externallyduring similar welding process can greatly reduce the longitudi-nal tensile RS in FSW joints. In our case, the effectiveness of theirmethod was investigated for RS mitigation in similar weld and

H. Eisazadeh et al. / Journal of Materials Processing Technology 222 (2015) 344–355 351

Fig. 16. Effect of thermal properties on formation of RS by making only one set of property dissimilar at a time. Case 1: similar weld; Case 2: thermal conductivity; Case 3:specific heat; Case 4: yield strength; Case5: thermal expansion; Case 6: fully dissimilar weld.

F1

3aot

F(

ig. 17. Longitudinal stress distribution during similar welding of 304 with a00 MPa tensioning load.

04-A36 dissimilar weld, respectively. At first, a similar weld of 304 plate was considered. A constant tensioning load in termsf a uniform longitudinal tensile stress of 100 MPa was appliedhroughout the welding thermal cycle in both weld plates. The load

ig. 19. Development of longitudinal RS during welding and cooling at a location in the Wpoint P shown in Fig. 2).

Fig. 18. Longitudinal RS distribution after similar welding of 304 with a 100 MPatensioning load.

was eventually removed once the sample was completely cooleddown to the room temperature. The distribution of the longitudi-

nal stress at an instant during welding and the longitudinal RS aftercooling are as shown in Figs. 17 and 18, respectively. When com-pared to the results obtained for the case without tensioning load

Z subjected to a 100 MPa tensioning load at an arbitrary location in the weld path

352 H. Eisazadeh et al. / Journal of Materials Processing Technology 222 (2015) 344–355

-tensi

(ss

(t1sdItts

Ftt

Fig. 20. Longitudinal RS after cooling for un

Fig. 7), a significant reduction of about 100 MPa in the tensile RStress in the WZ region, at the expense of some increase in tensiletress away from the WZ, is obtained.

In Fig. 19, the evolution of longitudinal stress at a given nodelocation P in Fig. 2), is displayed. Before the welding begins, longi-udinal stress at this location is the applied tensile stress, which is00 MPa in this case. However, as time progressed, and the heatource approached to the location P, tensile stress continued toecrease, eventually reaching a compressive stress of −200 MPa.

nterestingly, this maximum compressive stress was the same forhe case of un-tensioned sample (Fig. 8). As time progressed fur-her, the evolution of the longitudinal stress followed almost theame path, until the tensioning load is released at the end of the

ig. 21. Schematic illustrating the mechanism of RS mitigation using mechanical tensioninhermal stresses after cooling, (c) longitudinal thermal stresses during welding in the preo the releasing of the tensioning load, and (e) longitudinal thermal stresses after releasin

oned and tensioned case, similar 304 weld.

cooling cycle, i.e. after 2500 s. Upon releasing the tensioning load,the stress stored in the weld plate was reduced by an amount of thestress corresponding to the applied tensioning load. The reductionin the tensile stress in the WZ is further illustrated by compar-ing its distribution on the A–A′ cross-section in Fig. 20. Overall, theobtained results are consistent with those reported in the literaturefor similar welds (Chakravarti et al., 1990).

The mechanism of stress reduction through mechanical ten-sioning can be explained as follows: as the arc moves along the

weld path, material behind the torch expands, and developscompressive stress, due to the constraining of its free thermalexpansion by the surrounding material (Fig. 21a). As the materialreaches its yield strength, which is very low at high temperature,

g in similar weld: (a) longitudinal thermal stresses during welding, (b) longitudinalsence of tensioning load, (d) longitudinal thermal stresses after welding and priorg the tensioning load.

H. Eisazadeh et al. / Journal of Materials Processing Technology 222 (2015) 344–355 353

F nt timc f the rt

icmeetissssbdbt

ig. 22. Predicted longitudinal stress during welding and cooling phases at differeircle: tensioned sample. Left half is A36 and right half is 304. (For interpretation ohis article.)

t yields plastically producing a large plastic strain. As the materialools down to the room temperature and the yield strength of theaterial increases, a significant amount of elastic (misfit) strain

volves due to this nonhomogeneous plastic strain in the materialven when all the thermal strains are relieved. This explainshe large tensile RS distribution in the WZ with the maximumn the welding direction, balanced by the far field compressivetress (Fig. 21b). When the tensioning load is applied to the weldamples, as shown in Fig. 21c, the surrounding materials are beingtretched elastically. The behavior of material behind the torch isimilar to the un-tensioned sample, with the tensile stress limited

y the yield strength. As a result, the resulting ellipsoidal stressistribution that is present after cooling to the room temperature,ut before the removal of the tensioning load, is slightly largerhan that of the un-tensioned sample (Fig. 21d). Unloading the

Fig. 23. Longitudinal RS after cooling for un-tensioned

es in the dissimilar weld of A36 and 304. Red square: un-tensioned sample, blueeferences to color in this figure legend, the reader is referred to the web version of

superimposed tensioning load on the weld RS profile, therefore,reduces the stress field uniformly by an amount equal to the tensilestress corresponding to the tensioning load (Fig. 21e).

For RS reduction in dissimilar weld, only tensioning load wasapplied to the plate which experiences highest RS, (304 steel for thisstudy). The reason is that the tensioning stress caused by tensioningload will combine with the RS that is developing in the weld as itcools. Thus this will increase the tensile RS by 100 MPa (amount ofstress caused by tensioning load). This will mean that, there willbe a much greater overlap between RS and their correspondingyield stress limit (Price et al., 2007). Therefore, if the developing RS

profile was not of a sufficient magnitude to cause plastic yieldingin un-tensioned weld plate, but it will in tensioned weld plate. Indissimilar weld, the highest RS happens in 304 steel while RS in A36is no longer near yield limit. If a uniform tensioning load is applied

and tensioned cases in dissimilar A36-304 welds.

354 H. Eisazadeh et al. / Journal of Materials Proce

Fig. 24. Longitudinal stress distribution in dissimilar weld during welding with100 MPa tensioning load.

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ig. 25. Longitudinal RS distribution in dissimilar weld after cooling and the removalf 100 MPa tensioning load.

o both A36 and 304 weld plates, there might be a reduction inensile RS stress in the WZ region of A36, however it causes somencrease in tensile stress magnitude away from the WZ in A36. Toum up, since the RS field in dissimilar weld is asymmetric (Fig. 12),he introduced external tensioning load need not necessarily beniform or symmetric. However, it should be determined based onhe particular design requirements.

Noting that the maximum tensile RS in the 304 reached its yieldtrength, while the A36 remained significantly below the yieldtrength, a uniform tensile stress of 100 MPa applied only on the04 plate was considered as tensioning load in this study. Theormation of longitudinal RS in A36-304 dissimilar tensioned andn-tensioned welds are studied by capturing the evolution of longi-udinal stress distribution along line A–A′ (Fig. 14), for a few selectedimes (Fig. 22). The resulting RS, shown in Fig. 23, indicates a largeensile stress reduction in 304 side. A tensile RS reduction of about0% was also obtained in the A36 plate, which was not subjectedo the applied mechanical tensioning load.

The distribution of the longitudinal stress during welding (athe instance when the welding torch reaches the middle of thehole weld path) and after cooling is shown in Figs. 24 and 25,

espectively. When comparing these stress distributions to thoseorresponding to the un-tensioned case (Figs. 10 and 11, respec-ively), it can be concluded that the tensile RS in the WZ of aissimilar weld can be controlled as needed by choosing an appro-riate tensioning load.

. Conclusions

RS formation and the effect of the variation of an individual setf material properties on RS formation in a dissimilar weld werenvestigated using an FEA model. The effectiveness of mechanical

ssing Technology 222 (2015) 344–355

tensioning technique, particularly in the context of dissimilar weld,was also studied. The major findings of the study are summarizedbelow:

• Difference in the thermal expansion coefficient plays the primaryrole on RS formation in dissimilar welds of different ferrous alloys.

• Throughout the welding and cooling phases, 304 with rela-tively larger thermal expansion coefficient induced a tensileload on A36, while A36 induced a compressive load. Thus,due to yield, a larger longitudinal RS was formed in 304plate of the dissimilar weld when compared to the similar304 weld.

• Mechanical tensioning can be very effective in mitigating the RSin the WZ for both the similar and dissimilar GTA welds.

• The mechanism of RS reduction due to tensioning can beexplained based on the fact that the stress produced by the com-bined effect of both the plastic and elastic deformation cannotexceed the yield strength, which eventually allows for an elasticrecovery of the local stress upon its removal.

• Since the RS formation is asymmetric in a dissimilar weld, themechanical tensioning load to mitigate RS need not be symmetricor uniform, but should be introduced based on the specific designrequirements.

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