Effect of material properties and mechanical

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

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Journal of Materials Processing Technology journal homepage: www.elsevier.com/locate/jmatprotec

Effect of material properties and mechanical tensioning load on residual stress formation in GTA 304-A36 dissimilar weld H. Eisazadeh a , A. Achuthan a , J.A. Goldak b , D.K. Aidun a,∗ a b

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

a r t i c l e

i n f o

Article history: Received 29 September 2014 Received in revised form 13 March 2015 Accepted 14 March 2015 Available online 24 March 2015 Keywords: Residual stress Similar weld Dissimilar weld Mechanical tensioning Residual stress reduction

a b s t r a c t Using a finite element analysis (FEA) model, the residual stress (RS) formation in an autogenous GTA dissimilar weld between austenitic stainless steel (304) and low carbon steel (A36) are analyzed. The effect of material properties on RS formation was determined by first considering a similar weld of 304 plates, and then changing only a selected mechanical property of the 304 plate on one side of weld to that corresponding to an A36 plate. Enforcing one set of mechanical property to be different at a time helped to isolate the role of these individual properties on the RS formation in the dissimilar weld. The effect of mechanical tensioning on dissimilar welds is then investigated. Results show that the longitudinal RS in both the similar and dissimilar welds can be reduced in the weld zone (WZ) by an amount equal to the stress corresponding to the applied mechanical tensioning load, as the tensioning load is removed after cooling. The mechanism of RS formation in dissimilar weld, and its mitigation by mechanical tensioning are determined by comparing the longitudinal stress evolution on a cross-section of the dissimilar weld plates under the mechanically tensioned and free conditions. © 2015 Elsevier B.V. All rights reserved.

1. Introduction Welding is a reliable and efficient metal joining process widely used in the infrastructure and heavy equipment industry, such as in the construction of steel bridges, shipbuilding, and installation of large pipelines. High joint efficiency, simple setup, and low fabrication costs are the advantages of this joining process. However, due to the localized heating, and subsequent cooling during welding, highly non-uniform temperature distribution occurs across the weld and the base metal (BM), resulting in the formation of significant residual stress (RS) in the weldment. The RS formation in welds, being a thermo-mechanical phenomenon that depends on many factors, makes its quantitative prediction quite challenging. The principal factors that determine the RS formation in a welded structure weldment are shown in a Fishbone diagram in Fig. 1. Understanding the distribution of RS induced by welding is critical in many industrial applications to determine the crack growth behavior and predict failure. Several studies on the effect of RS on the failure of dissimilar weld joint have been reported. For instance, Suzuki et al. (2012) reported significant stress corrosion cracking (SCC) as a result of

∗ Corresponding author. Tel.: +1 315 268 6518. E-mail addresses: [email protected], [email protected] (H. Eisazadeh), [email protected] (D.K. Aidun). http://dx.doi.org/10.1016/j.jmatprotec.2015.03.021 0924-0136/© 2015 Elsevier B.V. All rights reserved.

RS in the dissimilar welds between ferritic steels and austenitic stainless steels, which is widely used in the oil and gas industry. In general, the experimental determination of RS in dissimilar welds is quite challenging when compared to the similar weld, due to the differences in the BM properties, especially thermal expansion coefficient and yield strength, producing a relatively complex distribution of RS. Lately, a number of studies have used numerical models based on finite element analysis (FEA) to predict RS in dissimilar welds. For example, Ranjbarnodeh et al. (2011) determined longitudinal RS in butt joints of dissimilar steels and compared them with the stresses in similar steel joints. Similarly, Katsareas and Youtsos (2005) developed a 2-D FEA model relying on a simple implementation of the material property mismatch for RS prediction in dissimilar metal pipe welds. Lee and Chang (2007) studied the effect of yield and tensile strengths on RS by employing different carbon steels under both similar and dissimilar butt weld conditions. Anawa and Olabi (2008) used FEA to determine the optimized process parameters and develop statistical models for the welding of stainless steel and low carbon steel using a CO2 continuous-wave laser. Deng et al. (2009) determined the RS in a dissimilar metal pipe joint considering cladding, buttering, and post-weld heat treatment (PWHT). Similarly, Lee and Chang (2011) predicted the axial and hoop RS produced in high strength carbon steel pipe weld using a FEA model by employing a sequentially coupled 3-D thermal and solid-state phase transformation during welding.

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Fig. 1. Fishbone diagram illustrates the effect of various parameters on RS formation.

RS mitigation techniques have received increased attention over the past two decades. Most of these techniques, however, are applicable to similar weld. Zhang et al. (2005) have discussed the adaptation and optimization of some of these techniques, such as annealing and peening processes, for its use on similar welded joints. There also exist specific welding methods such as in-process cooling and global mechanical tensioning (GMT), which are designed to work by influencing the thermal strain mechanism that drives RS formation in the weld. Altenkirch et al. (2008) experimentally investigated the relationship between the final RS stress, distortion, and the level of GMT applied in situ during welding. The mechanical tensioning technique, where a tensile stress is introduced before welding and maintained throughout the welding and cooling phases, was also found to be very effective by Richards et al. (2008) in relieving the large longitudinal RS in similar welds. Wen et al. (2010) demonstrated that rolling the weld as an effective way to improve the mechanical properties, and minimize distortion. Studies with regard to mitigating RS in dissimilar welds are limited. Sedek et al. (2003) demonstrated that the thermal stress relieving techniques, such as the furnace annealing of welded joints, were not only ineffective for mitigating RS in dissimilar steel welds, but actually increased RS as well. This is due to the considerable difference in the thermal expansion coefficient of the joined steels, producing large misfit stresses. Hurrell et al. (2006) discussed a number of mechanical mitigation techniques in a review article. Broadly, they classified the mechanical mitigation techniques under three main categories: (a) surface mechanical treatment to induce compressive skin stress, (b) mechanical stress relief through thickness, and (c) weld design optimization to produce low/favorable RS levels and minimize distortion. Among all the techniques in these three categories, Song et al. (2010) identified weld overlay technique as the most effective for protecting a structural dissimilar metal weld. Kim et al. (2009) found that both the longitudinal and the transverse stress components decreased as the number of layers increased. The effectiveness of mechanical tensioning technique has not been studied on a dissimilar weld where the RS is not the same on the two base plates. In summary, it can be stated that RS in dissimilar weld is asymmetric and its highest magnitude occurs in the plate which

possesses greater yield strength and thermal expansion. In case of ferritic steel and austenitic steel dissimilar weld, the latter undergoes higher RS. The belief behind this is that materials cannot hold RS beyond their yield point. Once stress magnitude at any location of weld plate reaches to its corresponding yield point, that particular location starts deforming plastically. Since austenitic steel has greater yield point, the magnitude of RS in this metal will be larger. Thermal expansion of austenitic steel has a big role to play in this development, as well. Since this property is again greater in austenitic stainless steel, area undergoing tensile RS will be larger. Overall, even though in the past decades, significant amount of progress on RS modeling was made, the mechanism of various mitigation techniques, mechanical tensioning in particular, in terms of its dependence on material properties of the two plates, is not well understood. In the present study, we first used a FEA model to track the evolution of stress field at various locations during the welding and cooling periods, providing a better understanding of the mechanism of RS formation in dissimilar welds, particularly the influence of the difference in various material properties. Then the mechanism of the influence of tensioning load on RS reduction in similar and dissimilar welds was determined.

2. Methodology In the experiment, the GTAW process without filler metal was considered for three butt-welded joints of type 304 stainless steel and A36 low carbon steel. The welding parameters used for the simulations conducted in this work are shown in Table 1. The chemical compositions of 304 and A36 plates are provided in Table 2. The temperature fields and the evolution of the RS were investigated by means of a sequentially coupled thermo-mechanical formulation available in the ABAQUS commercial package. The details of the model are shown in Fig. 2. A weld plate of dimensions 72 mm × 50 mm × 5 mm was considered. Linear 8-node brick elements, with relatively finer elements in the 5 mm region on both sides of the weld path, were used for the model. Simulation of welding was realized by introducing a moving volumetric heat source region into the plate at the desired speed in the desired direction.

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Table 1 Welding 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 2 Chemical compositions (wt%) of A36 and 304 steels. Grade

Carbon

Manganese

Silicon

Phosphorus

Sulfur

Chromium

Nickel

Balance

A36 304

0.08–0.29 Max 0.08

0.40–1.20 Max 2

0.15–0.40 Max 1

0.04 Max. Max 0.045

0.05 Max. Max 0.03

None 18–20

None 8–10

Fe Fe

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

Fig. 2. Details of the plate geometry, mesh, and the thermal and mechanical boundary conditions. T1, T2 and T3 represent the location of the thermocouples. The location P and cross-section A–A were used to study the temperature and stress evolution.

h [W/m2 K]

T − To [K]

1.85 9.079 18.5 52.6

56 278 556 2778

temperature-dependent Young’s modulus and Poisson’s ratio. The thermal strain was computed using the temperature-dependent coefficient of thermal expansion. For the plastic strain, a rateindependent plastic constitutive equation is considered with the von Mises yield criterion, temperature-dependent mechanical properties and linear isotropic hardening rule. Since, the effect of phase transformation for low carbon steel on welding deformation is insignificant and the phase transformation does not occur in the austenitic stainless steel, it was ignored in the computational model (Deng et al., 2003).

2.1. Transient thermal analysis 2.3. Boundary conditions The heat from the moving welding arc is applied as a volumetric heat source with a double ellipsoidal distribution and is expressed by the following equations (Goldak et al., 1984). The ellipsoidal heat source distribution ahead and behind of the welding torch (arc) is given, respectively, as: √ 6 3ff QW −3x2 /a2 −3y2 /b2 −3z2 /c2 1e Q (x , y , z  , t) = e (1) √ e a1 bc  √ 6 3fr QW −3x2 /a2 −3y2 /b2 −3z2 /c2 2e Q (x , y , z  , t) = e (2) √ e a2 bc  where x , y and z are the local coordinates of the double ellipsoid model, relative to the torch location and aligned with the welded plates. The ff and fr are the fraction of heat deposited in the front and the rear parts, respectively. Note that ff + fr = 2.0. In this study, ff is assumed to be 1.4 and fr to be 0.6, in order to introduce the expected steeper temperature gradient in the ellipsoid, ahead of the torch. Qw is the power of the welding heat source, which can be determined from the welding current, the arc voltage and the arc efficiency. The arc efficiency  is taken as 70% for the GTA welding process (Deng and Murakawa, 2006). The parameters a1 , a2 , b and c, define the characteristics of the welding heat source, and can be adjusted to create a desired melted zone according to the welding conditions. The moving heat source is implemented using the DFLUX user subroutine tool in ABAQUS. 2.2. Mechanical analysis The strain induced by welding process can be divided into elastic, plastic, thermal and transformation strain. The elastic stress and strain was modeled using the isotropic Hooke’s law with

The initial temperature of the material prior to welding was set to 298 ◦ K. Thermal boundary conditions consist of the application of convective 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 surface temperature (Table 3). The metal surface in the WZ and the surrounding region, being at high temperature, is dominated by heat losses due to radiation, while losses due to convection dominate in the relatively lower temperature surfaces, away from the weld zone (Deng and Murakawa, 2006). Table 3 presents the temperaturedependent convection coefficients used in our study. A value of 800 W/m2 K [4–5] was considered for surfaces which are in contact with 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 without constraining 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 material properties: thermal conductivity (k), specific heat capacity (C), mass density (), latent heat of melting and latent heat of solidification. The temperature dependencies of these properties for both 304 and A36 are shown in Fig. 3a. The latent heat of melting and solidification is assumed to vary linearly between the solidus and the liquidus temperatures. For 304, the latent heat, the solidus temperature and the liquidus temperature are 261 kJ, 1673 ◦ K and 1727 ◦ K, respectively. Likewise, for A36, it is 247 kJ, 1738 ◦ K and 1817 ◦ K, respectively.

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Fig. 4. Comparison between the measured and simulated temperatures for the similar welding of 304.

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

Fig. 3. Properties of A36 and 304: (a) thermal properties, and (b) mechanical properties.

Mechanical properties of 304 and A36, which are temperature dependent, are shown in Fig. 3b. The elastic–plastic material constitutive model is implemented in the form of temperature dependent stress–strain curves. 3. Results and discussion 3.1. Experimental validation of the model The FEA model was validated by comparing the predicted temperature with experimental data obtained for the similar weld with 304 steel. The temperature histories at three different locations were measured using K-type thermocouples spot welded to the upper and lower surfaces of the weld plate (Fig. 2). The thermocouples T1, T2, and T3 were placed at 3, 5 and 8 mm away from the weld line (boundary), respectively. The temperature data was acquired at a rate of 10 readings per second. For the computational model validation, the temperature histories predicted by the computational model throughout the heating and cooling stages were compared to the experimentally measured values (Fig. 4). Overall, the model predictions agree well with the measurements. Near the peak, agreement was relatively stronger for locations T2 and T3, when compared to location T1.

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

This could be attributed to the difference in the proximity of these locations to the fusion boundary (FB). Being the nearest to the FB, T1 experiences the largest temperature change and the associated property changes, thereby being the most sensitive to the underlying assumptions and input properties. 3.2. RS formation Figs. 5 and 6 present the typical temperature and longitudinal stress distribution, respectively, on the surface of the plate during welding at the instance when the welding torch (arc) reaches the

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Fig. 7. Longitudinal stress in 304 similar weld after cooling (longitudinal RS). Fig. 9. Temperature distribution at the instance when the welding torch reaches the middle of the weld path in A36-304 dissimilar weld.

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

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

middle of the whole weld path. The temperature and longitudinal stress distribution are symmetric with respect to the weld path, as expected. The temperature in the vicinity of the torch is very high and, therefore, the material volume in this region is in a state of low yield strength and, endures low stress. A large amount of compressive stress is generated just ahead of the torch because of the severe temperature gradient in this region, which is in the heating phase. Severe temperature gradient also means the strength of the material increased quickly to room temperature value within a short distance away from the weld zone (WZ). In the region behind the weld torch (green region in Fig. 5), which is in the cooling phase, the temperature gradient gradually decreases as the temperature around the WZ continues to decrease, and it approached relatively more uniform temperature distribution. As the cooling progresses, the fusion zone material behind the torch contracts, initiating formation of a tensile RS. As the cooling completes, the tensile longitudinal RS grows to a larger region throughout the WZ, as shown in Fig. 7. The temperature and the stress history at an arbitrarily chosen location in the WZ and in the middle of the weld path (location P shown in Fig. 2) are plotted in Fig. 8. During the initial heat up prior to the instance at which the center of the welding torch passes P (0–10 s period), the severe localized temperature gradients in the vicinity of the weld line created compressive stress in this region. As the welding torch passed P, and the material around P started to cool down, the large compressive stress transformed into a large tensile 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 weld when the torch reached the center of the weld path. As expected, the temperature distribution was asymmetric due to the difference in thermo-mechanical properties of the two plates. Heat diffusion during the heating and cooling phases was relatively larger in A36 due to its higher thermal diffusivity when compared to the 304. A36, with higher thermal conductivity than 304, also experienced a relatively faster temperature rise than 304. The longitudinal stress distributions at this instance and after complete cooling are shown in Figs. 9 and 10, respectively. The stresses are higher in the 304 plate due to the relatively larger yield strength and thermal expansion 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

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far from line A–A (Fig. 14). As the welding torch reaches the location corresponding to the 8 s instance from the start, compressive longitudinal stress began to evolve, but mostly in the region close to WZ. As the welding torch traveled forward and reached A–A , and the temperature increased, a W shaped stress profile with a relatively low tensile stress in the WZ, large compressive stress next to the WZ, and large tensile stress away from the WZ, developed for both the similar and dissimilar welds. Subsequently, as the welding torch crossed the section A–A and continued to travel forward, the tensile stress in the WZ increased in magnitude. The compressive stress magnitude remained the same, but spread into the outer tensile region. The evolution of stress during the cooling stage, where thermal mismatch and the associated stresses gradually changed the non-uniform plastic deformation, eventually resulted in a drastically different stress profile(@2500 s). The outer region (±20 mm) with substantial tensile stress completely changed to large compressive stress while the tensile stress in the WZ increased significantly, reaching the yield strength in 304. Fig. 12. Longitudinal RS after cooling in similar A36, similar 304, and their combination for dissimilar weld, which is shown by blue color, (left half is A36 and right half is 304).

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

weld), A36 (similar weld) and their combination (dissimilar weld). As discussed above, for the same heat input and BM dimensions, the magnitude of RS in the dissimilar weld around the WZ (±7 mm) is relatively higher in 304 than that in A36 due to the fairly larger yield strength and thermal expansion of 304 when compared to A36. The RS profiles of A36 and 304 similar welds, when compared to that of their dissimilar weld (A36-304 weld), show a significant difference in the longitudinal RS magnitude, as shown in Fig. 12. The maximum value the longitudinal RS in 304 of the dissimilar weld is 50 MPa more than that of the 304 similar weld, while the maximum longitudinal RS in the A36 is significantly lower than the A36 similar weld. The longitudinal RS normalized with its yield strength (YS) is presented in Fig. 13 for the dissimilar weld, which shows that the RS in 304 reached its yield strength in the WZ. The formation of longitudinal RS in similar welds of 304, A36, and the dissimilar welds of 304-A36 plates, is studied by capturing the evolution of longitudinal stress distribution along line A–A , for a few selected times (Fig. 14). After 3 s from the start (location shown as 3 s in Fig. 14b), the longitudinal stress is very low due to the low temperature gradient since the welding torch (arc) is relatively

3.3. Sensitivity of RS to various material properties A sensitivity study was performed to identify the impact of the dissimilarity in various material properties of 304 and A36, namely thermal conductivity, specific heat capacity, yield strength, and coefficient of thermal expansion, on the formation of RS. This was achieved by characterizing the RS formation in the WZ of a 304 plate, with the particular property under consideration changed to the corresponding A36 property, but only for material on one side of the weld (referred to as A36 plate in this section, as shown in Fig. 15). This approach, where only a single material property was dissimilar across the weld at a time, permitted to study the impact of individual material properties qualitatively (not quantitative because of elasto-plastic constitutive model which is nonlinear and hence history dependent) without being influenced by the dissimilarity of other material properties. The results are summarized in Fig. 16. In case 1, since material properties are exactly the same on both sides of the weld, longitudinal RS is symmetric with respect to the weld path, as discussed earlier. In case 2, bell shaped RS distribution shifted slightly toward the right due to the difference in thermal conductivity between the plates. A36 has a much higher thermal conductivity coefficient than that of 304 in the temperature range below 1000 K (Fig. 3). Since higher thermal conductivity leads to higher thermal diffusivity for a given mass density and specific heat capacity (˛ = k/Cp ), heat diffusion is faster in the A36. Therefore, the temperature distribution is relatively more uniform in this region when compared to the locations away from the weld path where temperature gradient is relatively large. In effect, the temperature across the plate becomes uniform more quickly in A36. Therefore, the lower temperature gradient in the A36 plate induced less tensile RS which was limited to a smaller region near the weld path. Since the specific heat capacity was similar for the A36 and 304 in the temperature range below 1000 K, the obtained stress field for case 3 was very similar to case 1. The yield strength was varied in case 4, resulting in the lowering of RS in A36, which has the lower yield strength. This can be attributed to the dependence of the RS on the plastic strains in the WZ, which essentially limits the maximum stress to the yield strength of the weld material. Compared to thermal conductivity, specific heat, and yield strength, the effect of thermal expansion coefficient showed the largest sensitivity to the RS in the WZ (case 5). Since the thermal expansion coefficient of A36 was small compared to the 304, the RS dropped to a remarkably low value in A36, and increased to a larger value in 304 reaching its yield strength. The results corresponding to the original dissimilar weld between A36 and 304, were also provided as case 6 in Fig. 16 for comparison purposes. Interestingly, cases 5 and 6 produced very similar longitudinal RS states,

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Fig. 14. (a) Predicted longitudinal RS profiles across the mid-plane of the weld at different times (red square for similar weld, blue circle for dissimilar weld with left half A36 and right half 304), (b) locations of welding heat source at different time. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)

807 K, the average difference of thermal expansion coefficient A36 and 304 is almost 11 × 10−6 . With this given difference, it can be concluded a larger thermal strain will be generated in the 304 side when the weld cools from 807 K to room temperature. Sedek et al. (2003) also stated the variation in the thermal expansion coefficient between ferritic–pearlitic and austenitic stainless steels in dissimilar weld caused RS differences in these two weld plates. Highest RS occurred in the latter because of its greater thermal expansion coefficient. They suggested these high stresses could be reduced considerably by using a transition material having a coefficient of thermal expansion between those of the two base metals. Fig. 15 illustrates that during weld thermal cycle, the 304 plate with large thermal expansion coefficient can produce tensioning load on A36 plate, while enduring compression load by itself. Fig. 15. Schematic illustrating how the variation in thermal expansion coefficient in a dissimilar weld can introduce mutual tension/compression loading conditions that influence the RS formation.

indicating that the coefficient of thermal expansion plays the most dominant role in the evolution of RS in the dissimilar welds of A36 and 304. Following the calculation shown by Deng et al. (2009), within the temperature range from room temperature (298 K) to

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 externally during similar welding process can greatly reduce the longitudinal tensile RS in FSW joints. In our case, the effectiveness of their method was investigated for RS mitigation in similar weld and

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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.

Fig. 17. Longitudinal stress distribution during similar welding of 304 with a 100 MPa tensioning load.

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

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

was eventually removed once the sample was completely cooled down to the room temperature. The distribution of the longitudinal stress at an instant during welding and the longitudinal RS after cooling are as shown in Figs. 17 and 18, respectively. When compared to the results obtained for the case without tensioning load

Fig. 19. Development of longitudinal RS during welding and cooling at a location in the WZ subjected to a 100 MPa tensioning load at an arbitrary location in the weld path (point P shown in Fig. 2).

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Fig. 20. Longitudinal RS after cooling for un-tensioned and tensioned case, similar 304 weld.

(Fig. 7), a significant reduction of about 100 MPa in the tensile RS stress in the WZ region, at the expense of some increase in tensile stress away from the WZ, is obtained. In Fig. 19, the evolution of longitudinal stress at a given node (location P in Fig. 2), is displayed. Before the welding begins, longitudinal stress at this location is the applied tensile stress, which is 100 MPa in this case. However, as time progressed, and the heat source approached to the location P, tensile stress continued to decrease, eventually reaching a compressive stress of −200 MPa. Interestingly, this maximum compressive stress was the same for the case of un-tensioned sample (Fig. 8). As time progressed further, the evolution of the longitudinal stress followed almost the same path, until the tensioning load is released at the end of the

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 the stress corresponding to the applied tensioning load. The reduction in the tensile stress in the WZ is further illustrated by comparing its distribution on the A–A cross-section in Fig. 20. Overall, the obtained results are consistent with those reported in the literature for similar welds (Chakravarti et al., 1990). The mechanism of stress reduction through mechanical tensioning can be explained as follows: as the arc moves along the weld path, material behind the torch expands, and develops compressive stress, due to the constraining of its free thermal expansion by the surrounding material (Fig. 21a). As the material reaches its yield strength, which is very low at high temperature,

Fig. 21. Schematic illustrating the mechanism of RS mitigation using mechanical tensioning in similar weld: (a) longitudinal thermal stresses during welding, (b) longitudinal thermal stresses after cooling, (c) longitudinal thermal stresses during welding in the presence of tensioning load, (d) longitudinal thermal stresses after welding and prior to the releasing of the tensioning load, and (e) longitudinal thermal stresses after releasing the tensioning load.

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Fig. 22. Predicted longitudinal stress during welding and cooling phases at different times in the dissimilar weld of A36 and 304. Red square: un-tensioned sample, blue circle: tensioned sample. Left half is A36 and right half is 304. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)

it yields plastically producing a large plastic strain. As the material cools down to the room temperature and the yield strength of the material increases, a significant amount of elastic (misfit) strain evolves due to this nonhomogeneous plastic strain in the material even when all the thermal strains are relieved. This explains the large tensile RS distribution in the WZ with the maximum in the welding direction, balanced by the far field compressive stress (Fig. 21b). When the tensioning load is applied to the weld samples, as shown in Fig. 21c, the surrounding materials are being stretched elastically. The behavior of material behind the torch is similar to the un-tensioned sample, with the tensile stress limited by the yield strength. As a result, the resulting ellipsoidal stress distribution that is present after cooling to the room temperature, but before the removal of the tensioning load, is slightly larger than that of the un-tensioned sample (Fig. 21d). Unloading the

superimposed tensioning load on the weld RS profile, therefore, reduces the stress field uniformly by an amount equal to the tensile stress corresponding to the tensioning load (Fig. 21e). For RS reduction in dissimilar weld, only tensioning load was applied to the plate which experiences highest RS, (304 steel for this study). The reason is that the tensioning stress caused by tensioning load will combine with the RS that is developing in the weld as it cools. Thus this will increase the tensile RS by 100 MPa (amount of stress caused by tensioning load). This will mean that, there will be a much greater overlap between RS and their corresponding yield stress limit (Price et al., 2007). Therefore, if the developing RS profile was not of a sufficient magnitude to cause plastic yielding in un-tensioned weld plate, but it will in tensioned weld plate. In dissimilar weld, the highest RS happens in 304 steel while RS in A36 is no longer near yield limit. If a uniform tensioning load is applied

Fig. 23. Longitudinal RS after cooling for un-tensioned and tensioned cases in dissimilar A36-304 welds.

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

Fig. 24. Longitudinal stress distribution in dissimilar weld during welding with 100 MPa tensioning load.

• Difference in the thermal expansion coefficient plays the primary role on RS formation in dissimilar welds of different ferrous alloys. • Throughout the welding and cooling phases, 304 with relatively larger thermal expansion coefficient induced a tensile load on A36, while A36 induced a compressive load. Thus, due to yield, a larger longitudinal RS was formed in 304 plate of the dissimilar weld when compared to the similar 304 weld. • Mechanical tensioning can be very effective in mitigating the RS in the WZ for both the similar and dissimilar GTA welds. • The mechanism of RS reduction due to tensioning can be explained based on the fact that the stress produced by the combined effect of both the plastic and elastic deformation cannot exceed the yield strength, which eventually allows for an elastic recovery of the local stress upon its removal. • Since the RS formation is asymmetric in a dissimilar weld, the mechanical tensioning load to mitigate RS need not be symmetric or uniform, but should be introduced based on the specific design requirements.

References

Fig. 25. Longitudinal RS distribution in dissimilar weld after cooling and the removal of 100 MPa tensioning load.

to both A36 and 304 weld plates, there might be a reduction in tensile RS stress in the WZ region of A36, however it causes some increase in tensile stress magnitude away from the WZ in A36. To sum up, since the RS field in dissimilar weld is asymmetric (Fig. 12), the introduced external tensioning load need not necessarily be uniform or symmetric. However, it should be determined based on the particular design requirements. Noting that the maximum tensile RS in the 304 reached its yield strength, while the A36 remained significantly below the yield strength, a uniform tensile stress of 100 MPa applied only on the 304 plate was considered as tensioning load in this study. The formation of longitudinal RS in A36-304 dissimilar tensioned and un-tensioned welds are studied by capturing the evolution of longitudinal stress distribution along line A–A (Fig. 14), for a few selected times (Fig. 22). The resulting RS, shown in Fig. 23, indicates a large tensile stress reduction in 304 side. A tensile RS reduction of about 10% was also obtained in the A36 plate, which was not subjected to the applied mechanical tensioning load. The distribution of the longitudinal stress during welding (at the instance when the welding torch reaches the middle of the whole weld path) and after cooling is shown in Figs. 24 and 25, respectively. When comparing these stress distributions to those corresponding to the un-tensioned case (Figs. 10 and 11, respectively), it can be concluded that the tensile RS in the WZ of a dissimilar weld can be controlled as needed by choosing an appropriate tensioning load. 4. Conclusions RS formation and the effect of the variation of an individual set of material properties on RS formation in a dissimilar weld were investigated using an FEA model. The effectiveness of mechanical

Altenkirch, J., Axel Steuwer, A., Peel, M.J., Withers, P.J., Williams, S.W., Poad, M., 2008. Mechanical tensioning of high-strength aluminum alloy friction stir welds. Metall. Mater. Trans. A 39 (13), 3246–3259. Anawa, E.M., Olabi, A.G., 2008. Control of welding residual stress for dissimilar laser welded materials. J. Mater. Process. Technol. 204 (1), 22–33. Attarha, M.J., Sattari-Far, I., 2011. Study on welding temperature distribution in thin welded plates through experimental measurements and finite element simulation. J. Mater. Process. Technol. 211 (4), 688–694. Biswas, P., Kumar, D.A., Mandal, N.R., Mahapatra, M.M., 2011. A study on the effect of welding sequence in fabrication of large stiffened plate panels. J. Mar. Sci. Appl. 10 (4), 429–436. Chakravarti, A.P., Malik, L.M., Goldak, J.A., 1990. Distortion Control in Simulated OnePass Panel Bead Welds. Deng, D., Murakawa, H., 2006. Numerical simulation of temperature field and residual stress in multi-pass welds in stainless steel pipe and comparison with experimental measurements. Comput. Mater. Sci. 37 (3), 269–277. Deng, D., Luo, Y., Serizawa, H., Shibahara, M., Murakawa, H., 2003. Numerical simulation of residual stress and deformation considering phase transformation effect. Trans. JWRI 32 (2), 325–333. Deng, D., Ogawa, K., Kiyoshima, S., Yanagida, N., Saito, K., 2009. Prediction of residual stresses in a dissimilar metal welded pipe with considering cladding, buttering and post weld heat treatment. Comput. Mater. Sci. 47 (2), 398–408. Goldak, J.A., Chakravarti, A., Bibby, M., 1984. A new finite element model for welding heat sources. Metall. Trans. B 15 (2), 299–305. Hurrell, P.R., Everett, D., Gregg, A., Bate, S., 2006. Review of residual stress mitigation methods for application in nuclear power plant. In: ASME 2006 Pressure Vessels and Piping/ICPVT-11 Conference, pp. 801–812. Katsareas, D.E., Youtsos, A.,2005. Residual stress prediction in dissimilar metal weld pipe joints using the finite element method. In: Materials Science Forum, vol. 490. Trans Tech Publ, pp. 53–61. Kim, K.S., Lee, H.J., Lee, B.S., Jung, I.C., Park, K.S., 2009. Residual stress analysis of an overlay weld and a repair weld on the dissimilar butt weld. Nucl. Eng. Des. 239 (12), 2771–2777. Lee, C.-H., Chang, K.-H., 2007. Numerical analysis of residual stresses in welds of similar or dissimilar steel weldments under superimposed tensile loads. Comput. Mater. Sci. 40 (4), 548–556. Lee, C.-H., Chang, K.-H., 2011. Prediction of residual stresses in high strength carbon steel pipe weld considering solid-state phase transformation effects. Comput. Struct. 89 (1), 256–265. Price, D.A., Williams, S.W., Wescott, A., Harrison, C.J.C., Rezai, A., Steuwer, A., Peel, M., Staron, P., Kocak, M., 2007. Distortion control in welding by mechanical tensioning. Sci. Technol. Weld. Join. 12 (7), 620–633. Ranjbarnodeh, E., Serajzadeh, S., Kokabi, A.H., Fischer, A., 2011. Effect of welding parameters on residual stresses in dissimilar joint of stainless steel to carbon steel. J. Mater. Sci. 46 (9), 3225–3232. Richards, D.G., Prangnell, P.B., Williams, S.W., Withers, P.J., 2008. Global mechanical tensioning for the management of residual stresses in welds. Mater. Sci. Eng. A 489 (1), 351–362. Sedek, P., Brozda, J., Wang, L., Withers, P.J., 2003. Residual stress relief in MAG welded joints of dissimilar steels. Int. J. Press. Vessels Piping 80 (10), 705–713.

H. Eisazadeh et al. / Journal of Materials Processing Technology 222 (2015) 344–355 Song, T.-K., Bae, H.-R., Kim, Y.-J., Lee, K.-S., 2010. Numerical investigation on welding residual stresses in a PWR pressurizer safety/relief nozzle. Fatigue Fract. Eng. Mater. Struct. 33 (11), 689–702. Suzuki, H., Katsuyama, J., Morii, Y., 2012. Comparison of residual stress distributions of similar and dissimilar thick butt-weld plates. J. Solid Mech. Mater. Eng. 6, 574–583.

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Wen, S.W., Colegrove, P.A., Williams, S.W., Morgan, S.A., Wescott, A., Poad, M., 2010. Rolling to control residual stress and distortion in friction stir welds. Sci. Technol. Weld. Join. 15 (6), 440–447. Zhang, J.X., Xue, Y., Gong, S.L., 2005. Residual welding stresses in laser beam and tungsten inert gas weldments of titanium alloy. Sci. Technol. Weld. Join. 10 (6), 643–646.