Original Article
Influence of the clamps configuration on residual stresses field in friction stir welding process
J Strain Analysis 2015, Vol. 50(4) 232–242 Ó IMechE 2015 Reprints and permissions: sagepub.co.uk/journalsPermissions.nav DOI: 10.1177/0309324715573361 sdj.sagepub.com
Caterina Casavola, Alberto Cazzato, Vincenzo Moramarco and Carmine Pappalettere
Abstract Friction stir welding is a joining process developed in 1991 by The Welding Institute. This welding technique is a solidstate joining process leading to joints with good mechanical performance and low residual stresses. In all welding techniques, the clamping systems have an important role in determining the quality of the welds and mechanical characteristics. Even more in friction stir welding, the position of the clamps plays a critical role because it is mainly a mechanical process with high forces involved. In this article, the correlation between the residual stress field and configurations of clamps has been established numerically. For this purpose an uncoupled thermo-mechanical finite element analysis has been carried out. The mechanical loads due to the tool have been also implemented into the model. The thermal and mechanical models have been validated on temperature field recorded by an infrared camera and residual stress field measured by X-ray diffraction analysis. The friction stir welding test was conducted on 6-mm-thick 5754 H111 aluminium alloy plates.
Keywords Friction stir welding, residual stress, temperature field, finite element method, infrared thermography
Date received: 7 November 2014; accepted: 23 January 2015
Introduction Friction stir welding (FSW) has been developed by The Welding Institute (TWI) in 1991, and nowadays, it has become an important welding technique to join materials that are difficult to weld by traditional fusion welding technology.1–3 This relevant characteristic is due to the nature of solid-state welding process of FSW. In fact, the material is not led to fusion, and the joint is the result of the rotation and movement along the welding line of the tool that causes softening of material due to frictional heat and the stirring of the same. The FSW technique is divided into four main steps: plunge, dwell, translation and exit. During the plunge phase, the tool penetrates into the material; subsequently, it is held in position for a few seconds while in rotation. This phase, called dwell time, has the aim to heat and soften the material before welding. Then, the tool moves along the joint line carrying out the weld. Finally, the tool is pulled out from the material. First industrial applications of the FSW were on aluminium alloys and light alloys that, in general, are difficult to weld by traditional technique.2 Nowadays, this technology is used
with magnesium, copper, nickel, metal matrix composites (MMC) and polymeric materials2–8 and also on titanium9 and steel.10 Furthermore, the possibility to couple the FSW with laser technique has been developed to improve the capability of the FSW.11 Several advantages distinguish the FSW from other traditional welding techniques; for example, it does not require protective atmospheres or filler materials, and surface cleaning is generally not necessary. Furthermore, it requires only 2.5% of the energy needed for a laser weld, and it is suitable for realizing various types of joints.2 These characteristics are of interest for many industries such as offshore platforms, shipbuilding, railway or automotive or aerospace industry. Anyway, although many industrial applications are in common Dipartimento di Meccanica Matematica e Management (DMMM), Politecnico di Bari, Bari, Italy Corresponding author: Caterina Casavola, Dipartimento di Meccanica Matematica e Management (DMMM), Politecnico di Bari, Viale Japigia 182 – 70126 Bari, Italy. Email:
[email protected]
Casavola et al. use,12 some aspects of FSW related to mechanical performance, fatigue behaviour and residual stress level should be investigated.11,13–15 Residual stresses have a fundamental role in welded structures because they affect the way to design the structures (e.g. the safety coefficients), fatigue life and corrosion resistance.16,17 Several articles have dealt with the residual stress distribution in FSW welds through experimental tests. Donne et al.18 have measured residual stress distribution on FSW 2024Al-T3 and 6013Al-T6 welds by different techniques. They have shown that longitudinal and transverse residual stresses have a ‘M’-like distribution across the weld. Furthermore, the longitudinal residual stresses are higher than the transverse ones regardless on tool rotation speed, pin diameter and traverse speed. Peel et al.13 have reported a measurement of residual stress in FSW welded AA5083 specimens by synchrotron X-ray method, which shows that the weld bead is in tension in both the longitudinal and transverse directions. Moreover, they have shown that the longitudinal stresses increase with the traverse speed increase. Sutton et al.19 have shown, according to Donne et al., that on 2024-T3 aluminium friction stir butt welds, the longitudinal stress is the largest tensile component, and that the transverse stress is about 70% of the longitudinal stress. With the purpose of a clearer understanding of the FSW process, in recent years, some researchers have studied the development of FSW numerical models. Furthermore, besides a deeper knowledge of FSW, these numerical simulations have also the aim to guide the development of the process through the research of optimal parameters minimizing the amount of trial and error. Several authors deal with heat generation in FSW process. Schmidt et al.20 have found an analytical model to describe the heat generation. Song and Kovacevic21 have modelled the heat transfer using the finite difference method. Khandkar et al.22 have made an uncoupled thermo-mechanical model for some aluminium alloys and 304L stainless steel based on torque input for calculating temperature and then residual stress. Chen and Kovacevic23 have developed a threedimensional numerical model based on a finite element method to study the thermal impact and evolution of the stresses in the weld. Many of these numerical models are only thermal or thermo-mechanical models in which the mechanical force that the tool exerts on the workpiece is not considered. These should be included into the thermo-mechanical model because FSW is mainly a mechanical process, and the forces involved in this type of work are relevant. Thermal expansion can produce high stresses, and also it can lead to yielding of the material if the structure is fully constrained. Therefore, it must have particular attention to the choice of the boundary conditions to avoid overconstraining the thermal expansion. Then, in this article, the correlation between the residual stress field, due to the FSW process, and configurations of
233 Table 1. Chemical composition Al5754. Al (%) 96.1–95.5
Si (%)
Fe (%)
Mn (%)
Mg (%)
0.4
0.4
0.5
2.6–3.2
Table 2. Physical and mechanical characteristics. Parameter
Value
Density (kg/m3) Melting temperature (°C) Thermal conductivity (W/m K) Thermal expansion coefficient (1/K) sy (MPa) su (MPa) Vickers Hardness HV
2650 600 138 23.9 3 1026 103.4 215 55
Table 3. Process parameters of FSW. Traverse speed (cm/min) 35
Dwell time (s)
Rotational speed (r/min)
Tilt angle (°)
Plunge force (N)
3
500
2
20,000
clamps will be established numerically. To this purpose, an uncoupled thermo-mechanical finite element analysis will be carried out. Also, the mechanical effect of the tool will be implemented. The thermal model, based on Schmidt et al.20 equation, will be validated experimentally on temperature field recorded by an infrared camera during the FSW process. The residual stress field behaviour, also, will be validated based on X-ray diffraction analysis measurement.
Material and methods The FSW process has been carried out on AA5754H111 specimen, whose chemical composition, thermophysical properties and mechanical characteristics at room temperature (20 °C) appear in Tables 1 and 2. This material has good weldability and excellent resistance to corrosion, and it is used in several industrial application such as shipbuilding, vehicle bodies, and welded chemical and nuclear structures. Table 3 sums up the parameters used in the FSW process. The specimen geometry is made up of two buttwelded plates of 200 mm length, 100 mm width and 6 mm thickness. A 4-kW machine with an Uddeholm QRO 90 Supreme tool steel has been used for FSW process. The tool (Figure 1) has a flat shoulder of 10.75 mm radius (Rshoulder), a threaded cylindrical pin with a radius of 3 mm (Rprobe) and 5.8 mm height (Hprobe). The FSW has been carried out along the welding line reported in Figure 2 that was perpendicular to the rolling direction of the AA5754 sheets. In the same figure, the clamping system has been represented. Temperature measurement, residual stresses analysis
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and then numerical simulation have been carried out for the retreating side of the FSW welded plate. Due to the layout of the FSW machine, it was not possible to acquire the temperatures of both the advancing and retreating side.
emissivity e = 0.95 has been set on the camera. Moreover, the central zone of the specimen has not been painted to avoid paint inclusion into the welded joint.
Residual stresses measurement Temperature measurement The infrared camera NEC H2640 (range: 0 °C–2000 °C, resolution: 0.06 °C, accuracy: 62 °C or 62%, spectral range: 8–13 mm) has been used to acquire the temperature history and field during the weld process. The setup of the thermal camera has been reported in Figure 3. The angle between the camera and FSW tool axis has been set to 30°. Due to reflection problems and low emissivity of aluminium, the specimen has been painted with matte black acrylic spray paint. An
Hprobe
The residual stress measurement is the relevant task for the characterization of a welded structure. Several techniques have been developed during time,24,25 but just some of these are quantitative and non-destructive. One of these and probably the most used technique is X-ray diffraction that can measure the residual stress on the surface of the structures without damaging. In this article, residual stresses measurements have been carried out using a Xstress 3000 G3R Stresstech X-ray diffractometer. A Cr tube (l = 0.2291 nm) powered with a 30 kV, 8 mA current and a 3-mm collimator has been used. A 2Y diffraction angle of 156.7° has been selected for the measures. The residual stresses, both longitudinal (x-axis direction) and transverse (y-axis direction), have been measured along the centre line of the specimens, as shown in Figure 2. In order to increase the quality of the measure, the stress in each point has been measured at five different angles (0°, 622.5°, 645°) with a F oscillation of 63°. The detection distance has been set to 75 mm. Moreover, the paint has been removed with a solvent before measuring the residual stress to eliminate the effect of this on X-ray diffraction measurements.
Rprobe Rshoulder
Numerical model An uncoupled FE model has been carried out to calculate residual stress field. At first, a thermal analysis has
Figure 1. Tool geometry scheme.
200 mm
200 mm
Centre line
Advancing Side
X
Welding direcon
Weld line
43 mm
20 mm
4 mm 10 mm
Y
Retreang Side
Figure 2. Specimen and clamping system scheme. Blue areas represent the experimental constraints.
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been conducted to calculate the temperature history due to the welding process. Then, the calculated temperature field has been used as thermal input to the mechanical model to predict the residual stresses and strains. For this simulation, the finite elements commercial code ANSYS 14.5 has been used. With the aim of an accurate simulation of residual thermal stresses, temperature-dependent thermo-physical and thermomechanical properties have been used. Due to lack of literature data, the density of the material has been kept constant with the temperature. The parameters of the material have been summed up in Table 4.26 Because the FSW is a solid-state welding process, it has been assumed that there is no material melting. Consequently, enthalpy values are not considered in this model. Moreover, a multi-linear isotropic hardening has been adopted to describe the behaviour of the material. A three-dimensional transient thermal analysis has been developed to simulate the temperature history and field. Due to the geometrical symmetry of the problem and in order to decrease the elements number, only a half plate has been modelled. The thermal model has been meshed using 6000 SOLID90 elements.27 A thicker mesh has been used near welding line to describe more accurately the thermal behaviour where the temperature gradient is higher (Figure 4). In the model, the
30°
FSW machine
Figure 3. Experimental setup.
natural convection between aluminium and air has been set to 20 W/m2 °C on the top surface and on the lateral side of the specimen. The conduction between the specimen and the backing plate have been simplified, according to Song and Kovacevic21 and Khandkar et al.,28 introducing an artificially high convection coefficient of 300 W/m2 °C on the bottom face of the plate. This value has been chosen in the way to match the maximum temperature of the numerical model with experimental data. Finally, according to Chao et al.,29 the radiation heat loss has been neglected because temperature hardly exceeds 500 °C. According to Schmidt et al.,20 the welding process has been modelled as a moving thermal source along the welded zone with a radius equal to Rshoulder. The analytical expression used for simulate FSW tool heat generation is Q=
2 Fn mv 2 ((R3shoulder R3probe )(1 + tan a) 3 Rshoulder + R3probe + 3R2probe Hprobe Þ
ð1Þ
Equation (1) considers the tool geometry such as Rshoulder, Rprobe, Hprobe and a, where a is the cone angle of tool shoulder. The other parameters are related to FSW process: m and Fn are the friction coefficient and the normal plunge force, respectively, while v is the rotational speed (Table 3). The friction coefficient m has been set equal to 0.3.20 In accordance with the experimental tests, the normal force Fn has been set to 20,000 N in the model. In the second step of the thermo-mechanical analysis, the thermal histories simulated by the FE thermal model have been inputted in the mechanical simulation to calculate the stresses. To this aim, the SOLID186 elements have been used instead of the SOLID90 elements but keeping the same mesh and load step size. Moreover, in this model, also the mechanical effect due to the compression force applied by the tool has been taken into account. To simulate this force, a uniform axial pressure distributed on tool area has been included in the FE mechanical model. Initially, the mechanical constrains have been set according to the experimental setup (blue area in Figure 2) constraining a 4 3 43 mm area for the left and right clamp. For the below clamp, a 4 3 45 mm
Table 4. Temperature-dependent material properties for AA5754-H111.26. Temperature (°C)
Density (kg/m3)
Specific heat (J/kg K)
Thermal conductivity (W/m K)
Thermal expansion (m/m °C)
Young’s modulus (GPa)
Poisson’s ratio
20 100 200 300 400 500
2650
891.6 950 988.4 1015.2 1045.6 1106.6
138 147.2 152.7 162.7 152.7 158.75
23.90 E-6 24.40 E-6 25.13 E-6 26.15 E-6 27.15 E-6 28.15 E-6
103.4 – 97.88 42.11 23.07 –
0.34 0.34 0.35 0.365 0.365 0.34
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Welded area Kside
Ktop = 20 W/m2K Kside = 20 W/m2K
Kboom = 300 W/m2K
Kside
Figure 4. FSW scheme of model mesh and convection coefficient.
100 mm 16 mm
A
50 mm
50 mm End
Middle
Start
Figure 5. Scheme of temperature measured points.
area has been constrained. Subsequently, the constrained area in the numerical models has been increased in the same manner for all three clamps (10 and 20 mm dotted lines in Figure 2) to evaluate the effect of boundary condition on residual stress field. The constraints on the top surface of the specimen block the movements in all directions. Moreover, the bottom of the plate has been constrained in z direction. Once the welding simulation is concluded, the constraints are gradually released and the residual stresses due to FSW process have been evaluated.
Results In order to obtain plausible stress values in the simulated cases with different constraints, the previously described model has been validated on both temperatures and stresses for the actual case. To reach this aim, the temperature field of the specimen has been measured during the FSW test by an infrared thermo-camera. The recorded values have been compared with the data calculated by the numerical model. To fully
validate the model, the temperature versus distance from welding line obtained by the numerical simulation has been compared with the experimental data in three different positions (Figure 5). The start position and the end position correspond to the initial and final welding phases, and they are about 50 mm from the right and left edges of the plate, respectively. Instead, the middle step is in the half of the specimen. For each position, the temperature has been measured at 16, 26, 36, 46, 56, 66, 80 and 100 mm from the welding line. It has not been possible to measure the temperature nearer the centre line for the presence of processing burrs that modifies the emissivity of the surface and so the temperature measured. In Figures 6–8, the graphs of the temperature versus distance from the weld line have been plotted for both numerical and experimental data. The results show that the temperature decreases moving away from the welding line and the maximum thermal gradient is in the area near the tool. Generally, the numerical data show a good agreement with the experimental measurements. This is confirmed by R2 value analysis defined as
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400 350
Experimental
Temperature (°C)
300
Numerical
250 200 150 100 50 0 0
20
40
60
80
100
120
Distance from the center line (mm) Figure 6. Numerical versus experimental data in the start phase.
400 350
Experimental
Temperature (°C)
300
Numerical
250 200 150 100 50 0 0
20
40
60
80
100
120
Distance from the center line (mm) Figure 7. Numerical versus experimental data in the middle phase.
R2 = 1
SSres SStot
ð2Þ
where SSres is the sum of the squares of the distances between experimental data and the numerical curve; instead, SStot is the sum of the square of the distances of experimental values from a horizontal line through the mean of all Y values. The R2 is 0.996 in the start step (Figure 6), 0.999 in the middle phase (Figure 7), and finally 0.987 in the end step (Figure 8). To obtain a consistent validation of the numerical model, the trend of temperature versus time for both numerical model and experimental data has been plotted (Figure 9). This graph shows the temperatures of the central point indicated by A in Figure 5. In general, there is a good agreement between the numerical model and the experimental data with some difference in the heating and cooling phase. The use of heat
convection instead of heat conduction to reproduce the presence of the backing plate and the assumption of a constant temperature for simulated backing plate could explain these discrepancies. In fact, during the welding process, the temperature of the backing plate rises, and this could decrease the cooling rate. Some authors21,28 , however, demonstrate that the use of convection to simulate the effect of backing plate introduces negligible errors and provides computational advantages. The maximum temperature reached in the point A is 323.3 °C. This value of temperature has been used to determine the convection coefficient of the simulated backing plate. In order to validate the numerical model on the basis of experimental data, the convection coefficient of the simulated backing plate has been chosen in the way to fit the numerical maximum temperature with the experimental data in the point A.
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400 350
Experimental
Temperature (°C)
300
Numerical
250 200 150 100 50 0 0
20
40
60
80
100
120
Distance from the center line (mm) Figure 8. Numerical versus experimental data in the end phase.
400 350
Experimental
Temperature [°C]
300
Numerical
250 200 150 100 50 0 0
20
40
60
80
100
Time [s] Figure 9. Temperature versus time trend of the middle point A.
In order to validate the mechanical part of the model, X-ray diffraction stress measures have been carried out. These superficial measures have been compared with the numerical model prediction. In Figure 10, the stress in the welding direction, that is, the longitudinal stress, has been reported and compared with the numerical values. The longitudinal stress shows a ‘M’-like distribution across the weld as Donne et al.18 have shown. Moreover, the maximum stress value in a FSW weld is located on the edge of the bead described in Figure 10 by the vertical grey line at a distance of 10.75 mm from the welding line. In Figure 11, the stress in the transverse direction has been reported and compared with numerical results. The trend of stress is roughly constant with a value of 60 MPa along all the transverse direction of the plate. Moreover, according to the observations reported by
Sutton et al.,19 the maximum transverse stress is 70% lower than the longitudinal stress. The distribution of the residual stress, both longitudinal and transverse, shows a good agreement with the experimental results. However, away from the welding line, there is maximum difference between numerical and experimental data for both longitudinal and transverse stresses. This could be explained considering that the initial residual stress in the numerical model is difficult to take into account. Moreover, away from the welding line, the influence of the welding process is minimal, and the residual stresses trend should tend to the pre-weld value. The numerical model has been solved changing the clamp positions as indicated in Figure 2. In Figures 12 and 13, the comparison of the residual stress profile between the actual clamp model and the two different
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Longitudinal stress
Longitudinal stress 250
200
Experimental
200
Numerical 4 mm
150
Numerical
150
Numerical 10 mm Numerical 20 mm
100
Stress [MPa]
Stress [MPa]
250
50 0
100 50 0
50
50
100
100
150 0
10
20
30
40
50
60
70
150
80
0
10
Distance from welding line [mm]
Figure 10. Numerical versus experimental comparison of longitudinal stress.
Transverse stress Experimental
200
Numerical 4 mm
Numerical
150
Numerical 10 mm Numerical 20 mm
100 50 0
100 50 0
50
50
100
100
150
80
Transverse stress
250
Stress [MPa]
Stress [MPa]
150
70
Figure 12. Numerical simulation of the clamp position on the longitudinal stress.
250 200
20 30 40 50 60 Distance from welding line [mm]
150
0
10
20 30 40 50 60 Distance from welding line [mm]
70
80
Figure 11. Numerical versus experimental comparison of transverse stress.
simulated clamp setups have been reported. The longitudinal stress (Figure 12) shows a reduction in the residual stress value in the welded zone and in the zone close to the bead. In fact, on the welding line, the longitudinal stress goes from 108 MPa of the actual setup to 83 MPa of 10 mm and 72 MPa of 20 mm setups. This reduction is more consistent from 4 to 10 mm than from 10 to 20 mm setups. Instead, away from the bead, there is the opposite behaviour, and 20-mm configuration shows an increase in residual stress of 30 MPa than the experimental setup. Figure 13 shows the trend of the transverse stress. There is a small reduction in the residual stress between the actual clamps setup and the simulated 10 and 20 mm. Moreover, the 10 and 20 mm setups do not show any substantial difference in the residual stress values between them. The maximum residual stress difference between experimental configuration and 20-mm simulated setup is 10.8 MPa. Finally, based on numerical results, the 20-mm clamp setup should reduce the residual stress in the welded zone and produce a beneficial effect on the welded specimen. Figure 14 shows the residual stresses distribution, both the longitudinal and transverse direction, in a
0
10
20
30
40
50
60
70
80
Distance from welding line [mm]
Figure 13. Numerical simulation of the clamp position on the transverse stress.
cross-section perpendicular to the welding direction and passing through the A point of Figure 5. The residual stress distribution before and after the release of the constraints is plotted. Before the release of the constraints, the stress values are always higher than after, both for the longitudinal and transverse directions. In general, for the longitudinal direction, the stress distribution is roughly constant in depth. Instead, the stress in the transverse direction is constant in depth only away from the bead, while near the welding zone, the stress values increase in depth. Finally, the trend showed in Figures 10 and 11 is confirmed also inside the plate.
Conclusion In this article, a three-dimensional thermo-mechanical model has been developed in order to have a better knowledge of the influence of the clamp position on residual stresses trend. This model includes the mechanical action of the shoulder and the thermo-mechanical characteristic of AA5754 at different temperatures. Thermographic analysis of FSW weld process has been carried out to validate the model thermally. The
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Figure 14. Predicted residual stress contour in the longitudinal and transverse directions in a cross-section perpendicular to the weld direction. The welding bead on the left of the dotted line.
thermal FE model can describe the heat-transfer process in FSW with a good correspondence with experimental data even if there are some discrepancies in the rising and cooling phase. These are mainly due to the assumption of heat convection instead of heat conduction to reproduce the presence of the backing plate. Moreover, the hypothesis of a constant temperature for
the simulated backing plate could contribute to explain these differences with experimental data. Furthermore, residual stress measurement by X-ray diffraction has been carried out to validate the FE mechanical part of the model. These experimental results confirm, as shown by Donne et al.,18 that the longitudinal stress has an ‘M’-like distribution across
Casavola et al. the weld, and the maximum stress value is located on the edge of the FSW bead. Instead, the transverse stress has lower values than longitudinal stress showing that this type of residual stress is a critical issue in FSW. The numerical model has been solved increasing the constrained area of the FSW plate. Following this change, the longitudinal stress shows a reduction in its value in the bead and in the zone close to the bead. Instead, away from the welding line, there is an opposite behaviour that increases the longitudinal residual stress bringing it closer to zero value. Moreover, the transverse residual stress, changing the clamped area on the plate, does not show a significant reduction in the residual stress. Finally, the numerical model shows that in the longitudinal direction, the stress distribution is roughly constant in depth. Instead, in the transverse direction, the stress is constant away from the bead, while it increases in depth near the welding zone. Further developments of the FSW FE model require the possibility to predict the difference between the advancing side and the retreating side for both the temperatures and residual stresses. These differences are due to the particular flow of the material during FSW process. Moreover, also the effect of the probe has to be considered in future works.
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Declaration of conflicting interests The authors declare that there is no conflict of interest. 16.
Funding This research received no specific grant from any funding agency in the public, commercial or not-for-profit sectors. 17.
References 1. Thomas WM, Nicholas ED, Needham JC, et al. Improvements relating to friction welding. Patent application PCT/GB92/02203, GB, 1991. 2. Mishraa RS and Ma ZY. Friction stir welding and processing. Mater Sci Eng R Rep 2005; 50: 1–78. 3. Casavola C, Lamberti L, Pappalettere C, et al. A comprehensive numerical stress – strain analysis of laser beam butt-welded titanium compared with austenitic steel joints. J Strain Anal Eng 2010; 45: 535–554. 4. Padmanaban G and Balasubramanian V. Selection of FSW tool pin profile, shoulder diameter and material for joining AZ31B magnesium alloy – an experimental approach. Mater Design 2009; 30: 2647–2656. 5. Lee WB and Jung SB. The joint properties of copper by friction stir welding. Mater Lett 2004; 58: 1041–1046. 6. Ferna´ndez JR and Ramirez AJ. Dissimilar friction stir welding of steel to Ni-based alloy 625 – butt and lap joints. In: Proceedings of the international offshore and polar engineering conference, Anchorage, AK, United State, 30 June–4 July 2013, pp.207–210. Cupertino (USA): ISOPE. 7. Cioffi F, Ferna´ndez R, Gesto D, et al. Friction stir welding of thick plates of aluminum alloy matrix composite
18.
19.
20.
21.
22.
23.
24.
with a high volume fraction of ceramic reinforcement. Compos Part A: Appl S 2013; 54: 117–123. Bagheri A, Azdast T and Doniavi A. An experimental study on mechanical properties of friction stir welded ABS sheets. Mater Design 2013; 43: 402–409. Farias A, Batalha GF, Prados EF, et al. Tool wear evaluations in friction stir processing of commercial titanium Ti–6Al–4V. Wear 2013; 302: 1327–1333. Khodir SA, Morisada Y, Ueji R, et al. Microstructures and mechanical properties evolution during friction stir welding of SK4 high carbon steel alloy. Mater Sci Eng A 2012; 558: 572–578. Campanelli SL, Casalino G, Casavola C, et al. Analysis and comparison of friction stir welding and laser assisted friction stir welding of aluminum alloy. Materials 2013; 6: 5923–5941. Gibson BT, Lammlein DH, Prater TJ, et al. Friction stir welding: process, automation, and control. J Manuf Process 2014; 16: 56–73. Peel M, Steuwer A, Preuss M, et al. Microstructure, mechanical properties and residual stresses as a function of welding speed in aluminium AA5083 friction stir welds. Acta Mater 2003; 51: 4791–4801. Ma YE, Xia ZC, Jiang RR, et al. Effect of welding parameters on mechanical and fatigue properties of friction stir welded 2198 T8 aluminum–lithium alloy joints. Eng Fract Mech 2013; 114: 1–11. Tra TH, Okazaki M and Suzuki K. Fatigue crack propagation behavior in friction stir welding of AA6063-T5: roles of residual stress and microstructure. Int J Fatigue 2012; 43: 23–29. Donne CD, Biallas G, Ghidini T, et al. Effect of weld imperfections and residual stresses on the fatigue crack propagation in friction stir welded joints. In: Second international symposium on friction stir welding, Gothenburg, Sweden, 26–28 June 2000. Abington Hall (UK): The Welding Institute TWI Zucchi F, Trabanelli G and Grassi V. Pitting and stress corrosion cracking resistance of friction stir welded AA 5083. Mater Corros 2001; 52: 853–859. Donne CD, Lima E, Wegener J, et al. Investigations on residual stresses in friction stir welds. In: Proceedings of the third international symposium on friction stir welding, Kobe, Japan, 27–28 September 2001. Abington Hall (UK): The Welding Institute TWI Sutton MA, Reynolds AP, Wang DQ, et al. A study of residual stresses and microstructure in 2024-T3 aluminum friction stir butt welds. J Eng Mater: T ASME 2002; 124: 215–221. Schmidt H, Hattel J and Wert J. An analytical model for the heat generation in friction stir welding. Model Simul Mater Sc 2004; 12: 143–157. Song M and Kovacevic R. Thermal modeling of friction stir welding in a moving coordinate system and its validation. Int J Mach Tool Manu 2003; 43: 605–615. Khandkar MZH, Khan JA, Reynolds AP, et al. Predicting residual thermal stresses in friction stir welded metals. J Mater Process Tech 2006; 174: 195–203. Chen CM and Kovacevic R. Finite element modeling of friction stir welding – thermal and thermomechanical analysis. Int J Mach Tool Manu 2003; 43: 1319–1326. Barile C, Casavola C, Pappalettera G, et al. Analysis of the effects of process parameters in residual stress
242 measurements on Titanium plates by HDM/ESPI. Measurement: J of the Int Measurement Conf 2014; 48: 220– 227. 25. Barile C, Casavola C, Pappalettera G, et al. Feasibility of local stress relaxation by laser annealing and X-ray measurement. Strain 2013; 49: 393–398. 26.Long X. Finite element analysis of residual stress generation during spot welding and its effect on fatigue behavior of spot welded joint. PhD Thesis, University of Missouri, Columbia, MO, 2005.
Journal of Strain Analysis 50(4) 27. ANSYS(R). ANSYS User’s Manual v. 14.5, Swanson Analysis Systems, Inc. 2012. 28. Khandkar MZH, Khan JA and Reynolds AP. Prediction of temperature distribution and thermal history during friction stir welding: input torque based model. Sci Technol Weld Joi 2003; 8: 165–174. 29. Chao YJ, Qi X and Tang W. Heat transfer in friction stir welding – experimental and numerical studies. J Manuf Sci E: T ASME 2003; 125: 138–145.
