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Welding residual stress reduction by scanning of a defocused beam
Journal of Materials Processing Technology 212 (2012) 19–26
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Welding residual stress reduction by scanning of a defocused beam F. Tölle a,∗ , A. Gumenyuk a , A. Backhaus b , S. Olschok b , M. Rethmeier a , U. Reisgen b a b
BAM Federal Institute for Materials Research and Testing, Unter den Eichen 87, 12205 Berlin, Germany Welding and Joining Institute, RWTH Aachen University, Pontstrasse 49, 52062 Aachen, Germany
a r t i c l e
i n f o
Article history: Received 6 May 2011 Received in revised form 26 July 2011 Accepted 31 July 2011 Available online 6 August 2011 Keywords: Residual stresses Stress reduction High energy beam welding Post-weld heat treatment Laser scanner optics
a b s t r a c t The residual stresses in narrow electron or laser beam welds with high stress gradients are decreased without any contact surfaces or additional equipment by applying the welding beam after welding in a defocused mode for heating the material regions in a certain distance from the weld on both sides. In case of electron beam application, the beam is positioned and focused by the electromagnetic coil with high frequency. In case of laser beam application a laser scanner optics enables fast positioning by an optomechanic beam deflection, while defocusing of the laser beam is obtained by increasing the distance between scanner optics and workpiece. Dependent on the component geometry and on the beam power different process parameters are used. The adjustable process parameters are the radius and the power of the defocused beam and the transversal and longitudinal distances between the welding and the defocused beam. The mechanism and the influence of the process parameters are investigated by FEM-simulation and a number of experiments on a ferritic steel S355J2+N with 5 mm thickness. FEMsimulation is used to reduce the matrix of process parameters for the experiments. The best experimental result shows a stress reduction of about 70%. © 2011 Elsevier B.V. All rights reserved.
1. Introduction In welding, a small material region is melted and cooled down with high temperature gradients. As described by Nitschke-Pagel and Dilger (2006) in their review of the causes of welding residual stresses for diverse welding procedures and various materials, the heat-affected material shrinks during cooling and causes stresses. Especially due to high energy welding processes like electron or laser beam welding these stresses can reach the temperature and phase dependent local yield strength in longitudinal weld direction. Stresses ranging at this level will be relaxed by plastic deformations as shown by Nitschke-Pagel and Dilger (2007a) in their review concerning welding residual stress assessment. The authors point out that residual stresses can influence the performance especially for high-strength materials. In consequence of stress relaxation or stress corrosion cracking these high stresses can limit service life depending on the tensile strength of materials as discussed by Nitschke-Pagel and Dilger (2007a) referring to a number of fatigue tests. Han et al. (2002) substantiated this influence of welding residual stresses on the fatigue strength as well as on the component life and developed a model for quantitative prediction of residual stress relaxation under cyclic load. The influence of high tensile residual
∗ Corresponding author. Tel.: +49 30 8104 3101; fax: +49 30 8104 1557. E-mail address: fl[email protected] (F. Tölle). 0924-0136/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.jmatprotec.2011.07.019
stresses on stress corrosion cracking was discussed by Sun (1993) and Brauss et al. (1998), too. A large number of processes for lowering the welding residual stresses is available like stress relief annealing as mentioned and described by Nitschke-Pagel and Dilger (2007b) or the static low-stress-no-distortion-technique as developed and discussed by Guan et al. (2006) which works through generating a specified temperature profile in the welded component using heating and cooling elements. But these methods are better applicable for wider welds such as TIG welds and for simple component geometries to provide the required contact surfaces for heating and cooling elements. They are also cost-intensive. The more flexible dynamically controlled low-stress-no-distortion-technique invented and described by Guan et al. (1994) uses a trailing cooling jet for generating the required temperature profile. Nevertheless, van der Aa et al. (2007) proved that this method is only effective for cooling jet travel speeds of less than 8 mm/s. For faster welding processes like beam welding with ordinary welding speeds higher than 10 mm/s the penetration effect of this cooling is minimized. Other mechanical methods of stress reduction yield lower tensile stresses only in near-surface areas. Water jet peening is one of such methods. Mochizuki (2007) showed that this peening effect only introduces compressive stresses in the treated surface and enhances the fatigue strength due to the reduction of stress corrosion cracking. Beam welding processes offer the advantages of low component deformations, low heat input and remote application, which enable joining of complex component geometries. The previously
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Fig. 1. Process scheme for simulation and experimental investigation of a welding beam and heating fields produced with the important geometric process parameters: radius of the heat source r, longitudinal distance between welding source and heating fields dx and transversal distance between welding source and heating fields dy .
designed stress reducing techniques are not practicable for this fast remote application because of the required contact surfaces. Therefore, a research concept for a stress mitigating process applicable to beam welded components is presented in this paper aimed at developing a contactless technique operating without any additional equipment. As shown by Rosen et al. (2007), application of the electron or the laser beam after welding for the reduction of high longitudinal residual stresses enables such a remote process. In this method the beam is used in a defocused mode for heating the material regions on both sides of the weld from a certain distance to generate additional tensile stress zones which unload the weld. Due to this unloading effect the residual stresses in the weld are decreased. The process scheme of welding and heat treatment in one work stage, i.e. simultaneous welding and heat treatment, is shown in Fig. 1. It is also possible to perform the heat treatment in a second work stage separated from the welding process. The important investigated process parameters, which will be explained in this article, are the defocused beam radius r, the maximum temperature Tpost,max to be generated by the defocused beam in the heated regions, and the transversal dy and longitudinal dx distances between the weld centre and the defocused beam. In Rosen’s (2007) work only the feasibility of this process for stress reduction was investigated without an analysis and any explanation of the stress reduction mechanism in this method. Furthermore, these researches are restricted for material thicknesses of 2 mm and 3 mm, a defocused beam radius of 10 mm (achieved stress reduction of about 85%) and only for the electron beam application. At this point the investigations in this article continue Rosen’s work. In these researches the stress reduction mechanism is described and the whole process parameters were varied and investigated. In addition, this method was applied to a laser beam welding source. The process parameters correlate with each other and generate a huge parameter matrix for the experimental investigations. To reduce the dimension of this matrix for the experiments with an electron and a laser beam, respectively, a qualitative parameter study was performed by FEM-simulation. With the simulative and the experimental results advices for the most suitable parameter sets can be concluded. 2. Finite element simulation to investigate the influence of the process parameters For the parameter study with the ANSYS software a simplified simulation model had to be created. This model was only a tool for a qualitative investigation of the stress reduction mechanism and to define process parameter regions for the experimental researches. Based on the resulting symmetry conditions (cf. Fig. 1)
Fig. 2. Scheme of the dependence of the power density distribution of an electron beam from the position of the beam focus to the workpiece.
in the finite element model, only half of the specimen was simulated to decrease simulation time. Another simplification was the use of a two-dimensional concept for the thermal and the structural model with defined z-depth characterizing the specimen thickness of 5 mm. Thus, the adiabatic thermal model represents ideal heat conduction in the heating fields caused by the defocused beam in thickness direction (z-direction). The adiabatic model was used because the two-dimensional model should represent a layer in the middle of the plates. In this model no heat transfer in vertical direction takes place and this corresponds to the adiabatic boundary conditions on the upper and the lower surfaces. The cooling of this layer due to heat conduction in thickness direction and due to the exchange with the environment is complex. So for this qualitative investigation, which parameter regions are most suitable for the stress reduction, it is not important to cool down the simulated plates to room temperature. A further cooling would result in an increase of the stresses but do not influence the most qualified parameter regions for the experimental investigations because the stress reducing mechanism is induced during the heat treatment and not during cooling down after this heat treatment. In the structural model the plane stress condition is illustrated by the two-dimensional layout. The elastoplastic material model was designed based on the temperature dependent material properties of S355J2+N from the SYSWELD database (ESI Group, 2006). To consider phase transformation during the welding process, a peak temperature driven approach as explained by Gebhardt et al. (2010) was used. This approach based on a change of the material properties if the element temperatures reach specific values. If they are larger than 740 ◦ C the martensitic material model for the yield strengths and other thermal expansions were used. The heat source for welding was modelled as a traveling point heat source and for the post welding heat treatment representing the defocused beam a cyclic area with a specified radius and a homogeneously distributed power density was modelled. This approach for the power density distribution correlates with measurements of a defocused electron beam in over focus position, as seen in Fig. 2. In Fig. 2 the dependence of the power density distribution of an electron beam from the position of the beam focus to the workpiece is shown. With the distribution in over focus position the smoothest temperature gradient in the heat treated areas is generated. So this over focus position was used for the experimental investigations with the electron beam welding source and for modelling the defocused heat source.
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Fig. 3. Simulated longitudinal stress distribution in the workpiece across to the welding direction after welding (solid line) and after applied heat treatment (dashed lines) with different temperatures Tpost,max generated by the defocused beam influencing the stress reduction in the weld with the process parameters: r = 15 mm, dy = 25 mm, dx = 200 mm and vhf = 15 mm/s.
This homogeneously distribution of the power density does not represent the distribution in a defocused laser beam. Therefore, the conclusions for the most suitable process parameters resulting from the simulation and the experiments with the electron beam were adapted to the experiments with the laser beam and checked for their applicability. With this FEM-model it is possible to investigate the qualitative influence of the process parameters. 3. Results of the finite element simulations Fig. 3 shows the influence of the temperature Tpost,max generated by the defocused beam on the stress reduction potential in the weld. This figure as well as Figs. 4 and 5 correspond to the heat treatment conducted in a second work stage (dx = weld length = 200 mm) separated from the welding process. During cooling, stresses in the range of around 600 MPa were generated in the weld (welding speed vw = 15 mm/s, beam power Pw = 800 W). The Tpost,max value was determined by the beam radius r, the travel speed vhf and the beam power Phf used for this heating field. Up to temperature of about 700 ◦ C, the stress reduction increases with rising temperatures. This behaviour can be explained by analysing the twodimensional stress distribution illustrated in Fig. 4 (with process parameters r = 10 mm, dy = 27 mm, dx = 200 mm, travel speed v = 5 mm/s and Tpost,max = 801 ◦ C) and the time dependent tem-
Fig. 4. Simulation result for the longitudinal stresses in MPa after heat treatment with the process parameters: r = 10 mm, dy = 27 mm, dx = 200 mm, vhf = 5 mm/s and Tpost,max = 801 ◦ C.
Fig. 5. Temperature (top) and longitudinal stress (bottom) profile over time in the weld (Tweld , weld ), in the centre of the heating field (THF , HF ) and in not heat treated region (TNHT , NHT ) with the process parameters: r = 10 mm, dy = 27 mm, dx = 200 mm, vhf = 5 mm/s and Tpost,max = 801 ◦ C (positions in the model are shown in Fig. 4; tbHT indicates the point in time of heat treatment, ttu indicates the point in time of termination of the unloading effect).
perature and stress profiles reflected in Fig. 5. In Fig. 4, the two-dimensional longitudinal stress profile is plotted after the heat treatment and a subsequent cooling time of 8000 s. In Fig. 5, the corresponding temperature profiles in the weld Tweld (cf. Fig. 4 coordinates x = 100 mm and y = 0 mm), in the heating fields THF (cf. Fig. 4 coordinates x = 100 mm and y = 27 mm) and in the non-influenced area TNHT (cf. Fig. 4 coordinates x = 100 mm and y = 84 mm) are correlated with the longitudinal stress profiles ( weld , HF , NHT ) in the same points. They were recorded over the time during processing on the simulated welded plate. After welding, high tensile longitudinal stresses weld were present in the weld. Due to heating by the defocused beam, the temperature dependent yield strength decreases in the heated region, and thus the generated compressive stresses (cf. Fig. 5 HF between 30 s and 35 s) can relax based on plastic deformations. The yield strength of low alloyed ferritic steels decreases with a high gradient in the temperature range between 600 ◦ C and 800 ◦ C during heating. If the generated temperature ranges below 700 ◦ C, the induced compressive stresses are lower and the yield strength in lower temperature regions is too high for stress relaxation. Temperatures above 800 ◦ C or 900 ◦ C give only marginal further yield strength reductions, i.e. the most suitable Tpost,max is about 700 ◦ C for the ferritic steel S355J2+N. While the heat treated regions cool down, the compressive stresses decrease and switch to tensile stresses due to material shrinkage. These additional tensile stress zones on both sides of the weld lead to a mechanical unloading of the weld as shown in Fig. 5 by weld in the time interval between 30 s and 70 s. The magnitude of the evolution of longitudinal stresses in the weld and of the unloading effect in dependence on the longitudinal distance dx between melting and post welding heat treatment processes can be seen in Fig. 6, too. This diagram shows the stress generation in the weld during cooling down from the melting temperature for an as welded plate compared to welded plates with a subsequent heat treatment at different dx while processing. The larger dx , the higher are the tensile stresses which can be generated in the cooling time interval (cf. Figs. 5 and 6 till the point in time before heat treatment tbHT ) between the welding and the reheating process. During cooling of the additionally heated regions down to
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Fig. 6. Evolution of longitudinal stresses in the weld dependent on the longitudinal distance between the welding source and the heating fields dx while cooling from melting temperature (tbHT indicates the point in time of heat treatment, ttu indicates the point in time of termination of the unloading effect); process parameters: r = 15 mm, dy = 25 mm, vhf = 15 mm/s and Tpost,max = 830 ◦ C.
the same temperature as that in the weld zone (cf. Figs. 5 and 6 for Tweld ≈ THF at the point in time of termination of unloading ttu = 60 s), the tensile stresses in the weld are decreased by unloading. In other words, the unloading occurs as long as the heat treated regions shrink more than the weld zone. As a result of heat conduction the weld is reheated. At the point in time ttu , the temperatures in the heated zones and in the weld are equal and the unloading terminates. It can also be seen in Fig. 6 that a shift to larger dx , i.e. longer cooling times before heat treatment tbHT , increases the magnitude of the unloading effect. Further cooling of the weld and the heat treated regions causes an increase in the longitudinal stresses. Due to the idealized conditions in the adiabatic two-dimensional model the workpieces exhibit a homogeneously distributed temperature of about 270 ◦ C at the end of the simulations after around 8000 s cooling time. For small dx (nearby 0 mm), the temperature in the weld is higher than in the additionally heat treated regions. In this case no mechanical unloading occurs through the heat treatment. If dx is so large that the weld is able to cool down to ambient temperature, a further increase of dx does not lead to significantly larger stress reductions. In further simulations, the influences of the transversal distance dy and of the defocused beam radius r on stress reduction were investigated. In this respect, the most suitable dy depending on the defocused beam radius was determined to obtain the maximal value of the stress reduction in the weld zone. Fig. 7 represents the maximum stresses in the weld depending on dy at different radii r. Each single point shows the value of the calculated longitudinal residual stress in the weld corresponding to a specific beam radius r and a certain transversal distance dy . From the diagram given in Fig. 7 it can be concluded that dy ≈ r + 12 mm provides the most suitable stress reduction of this process. With these simulation results the stress reduction mechanism can be illustrated and parameter regions which are most suitable for the experimental investigations can be defined. 4. Experimental For qualifying this process for stress reduction in the weld, a number of experiments with varied process parameters were carried out on 5 mm thick ferritic steel S355J2+N. Parameters resulting from the presented FEM-simulations were used. The applied workpieces were free from residual stresses before welding.
Fig. 7. Correlation of the radius of the heat source r with the transversal distance between welding source and heating fields dy (dx = 200 mm, vhf = 15 mm/s and Tpost,max ≈ 700 ◦ C), the longitudinal residual stresses in the weld centre (each point represents one simulation result) are shown (stress in the weld without additional heat treatment in grey at 580 MPa).
In the electron beam welding experiments, full penetration welds were produced by the focused beam with 0.3 mm beam radius and 6.12 kW beam power Pw . The welding speed vw was 15 mm/s for all experiments. Workpiece dimensions of 200 mm × 200 mm × 5 mm were used. Due to the feasibility of high frequency electromagnetic beam focusing and deflection in electron beam welding it is possible to use the beam for the heat treatment in the defocused mode simultaneously with welding. Therefore, the electron beam was positioned and deflected with a frequency of 500 Hz to the three positions in the following sequence: weld position, heat treatment position to the left of the weld, weld position, heat treatment position to the right of the weld, back to the weld position. For welding and heat treatment in one work stage the same speed for all three positions must be used (vw = vhf = 15 mm/s). Due to the machine dependent maximum deflection distance of 70 mm, experiments with different travel speeds for welding and heat treatment (vw = / vhf ) as well as experiments with longitudinal distances larger than 70 mm had to be performed in a second work stage. In the cases of two work stages, the focused beam was used in the first work stage for the welding process with the same welding parameters as applied to the in situ heat treatment process. In the second work stage the defocused beam was deflected between the two positions of heat treatment with a frequency of 500 Hz, too, and moved along the weld with a travel speed vhf of 15 mm/s. In these cases, the weld was allowed to cool down to temperatures below 200 ◦ C before the heating process was started in a second work stage. Under these conditions, the time variable tx describing the delay between the end of welding and the beginning of heat treatment was used instead of the longitudinal distance dx . The different beam radii for the heat treatment were generated by varying the focus current and controlled by measurements of the power density distribution. For three radii (6.5 mm, 10.0 mm and 15.0 mm) aligned powers Phf (2.88 kW, 3.12 kW and 5.28 kW) of the defocused beam were used for heating the workpiece surface. The parameter regions investigated in the electron beam experiments are listed in Table 1. The second part of the experimental program was carried out with laser beam welding using a scanner optics that operates by optomechanic beam focusing and deflection. However, as the deflection speed of the laser optics was limited to 14 Hz and was much slower than the electromagnetic deflection of the electron beam, it was not possible to conduct keyhole welding with simultaneous heat treatment at different positions in one work stage. In
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Table 1 Investigated parameter regions in the electron beam experiments. Radius of the defocused beam [mm]
dy [mm]
tx [s]
Travel speed vhf [mm/s]
Power for defocused beam Phf [kW]
6.5 10.0 15.0
0, 5, 10, 15, 20, 30 23 26
0–4.66, 18–43 18, 30 18, 30
2.5, 5, 15 5, 15 5, 15
0.48–2.52 1.25, 3.12 2.11, 5.28
this case, heat treatment in laser beam processes had to be performed in a second work stage. For these experiments a 5 kW fibre laser coupled with the laser scanner optics was used with the following welding parameters: 5 kW beam power Pw , 0.3 mm beam radius of the focused laser beam and 25 mm/s welding speed vw . To obtain different radii of the defocused beam, the beam waist was positioned in specific distances above the specimen surface. The three different radii 9.5 mm, 12.0 mm and 15.5 mm (including 86% of the beam power of the defocused beam) in the workpiece plane were controlled by measurements with a power density meter. After welding and cooling down to ambient temperature, the second work stage was executed by lifting the scanner optics to the adjusted distance and using different powers for the specific beam radius at a travel speed vhf of 5 mm/s. Accordingly, for laser beam experiments it was only possible to investigate the effect of this process when the temperature in the weld was low. This means that the variable tx adopted to approach infinity and was not specified for the individual experiments. The duration of 60 ms was used for heating one position before the beam was deflected to the other side of the weld for heating. With the help of a thermographic camera the necessary power values Phf for heating the workpiece surface in the heating field centre to about 700 ◦ C were aligned to obtain corresponding values for the defocused beam radius. Using X-ray diffraction, the residual stresses were measured in the workpiece surfaces (averaged over the X-ray penetration depth of up to 100 m) after the experiments. Calculation of the stresses was carried out adopting the sin2 -method as described by Spieß et al. (2005).
Fig. 8. Experimentally measured longitudinal stresses in an electron beam reference weld (welding speed vw = 15 mm/s, beam power Pw = 6.12 kW) on the top and bottom surface compared to a specimen welded with the same parameters after heat treatment with the following parameters: vhf = 15 mm/s, beam power for heat treatment Phf = 5.28 kW, defocused beam radius r = 15 mm, dy = 26 mm, time interval between welding source and heating fields tx = 30 s.
higher laser power levels and travel speeds for heating the material surface to this temperature region are not possible. Table 3 lists the best results of the X-ray diffraction measurements in the laser beam investigations showing similar stress reductions compared to the electron beam experiments. The longitudinal stress profile of the experiment with the best stress reducing result of 22% on the top and 73% on the bottom surface and the radius of 15.5 mm is compared to the reference weld in Fig. 9. The results listed in Tables 2 and 3 verify the simulation results concerning the influence and the values of the most suitable process parameters. In the experiments with both beam sources similar results were generated for the most suitable parameters. Accordingly, these results were compared together with the simulation results. As assumed from the FEM-simulations, in Tables 2 and 3 the lowest experimentally measured residual stresses in the weld for each radius are achieved with a dy value of about the radius plus 11–15 mm (hypothesis from simulation results dy ≈ r + 12 mm). In Figs. 8 and 9 the longitudinal stress profiles for the best experienced parameter sets are compared with reference welds for electron and laser beam welds. Remarkably, for both beam sources the stress
5. Experimental results and discussion of the stress reducing capability The best results obtained in the electron beam investigations are listed in Table 2. The results of the X-ray diffraction measurements reveal for the largest experienced radius of 15 mm a stress reduction of about 36% on the weld top surface and of 76% on its bottom surface. The longitudinal stress profile of this specimen with the best stress mitigating result is compared to the stress profile in a reference weld in Fig. 8. In the experiments with the laser beam, the desired temperatures in the heating fields of around 700 ◦ C for the defocused laser beam radius of 15.5 mm were reached with 4 kW beam power and 5 mm/s travel speed. Due to the maximum scanner capacity (5 kW)
Table 2 Longitudinal stresses in the electron beam weld centre reference weld without heat treatment. Radius of the defocused beam [mm]
Beam power for heat treatment Phf [kW]
vhf [mm/s]
dy [mm]
tx [s]
Longitudinal stresses in the weld on the top side [MPa]
Reference 6.5 6.5 6.5 6.5 6.5 6.5 10.0 15.0
– 2.88 2.88 2.88 0.96 1.12 1.44 3.12 5.28
– 15 15 15 5 5 5 15 15
– 10 20 35 20 20 20 23 26
– 5 5 5 5 5 5 30 30
471.4 430.3 384.7 500.0 468.3 374.8 440.2 377.0 304.1
± ± ± ± ± ± ± ± ±
13.0 32.3 60.6 54.1 63.3 45.1 61.9 58.3 53.8
Stress reduction on the top side [%]
Longitudinal stresses in the weld on the bottom side [MPa]
Stress reduction on the bottom side [%]
– 9 18 −6 1 20 7 20 36
375.1 ± 12.7 – 354.8 ± 23.0
– – 5
– – – 89.3 ±59.0 91.9 ±43.6
– – – 76 76
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Table 3 Longitudinal stresses in the middle of the laser beam weld (vhf = 5 mm/s). Radius of the defocused beam [mm]
Beam power for heat treatment Phf [kW]
dy [mm]
Longitudinal stresses in the weld on the top side [MPa]
As welded 9.5 9.5 9.5 12.0 12.0 12.0 15.5 15.5 15.5 15.5 15.5
– 2.3 2.3 2.3 3.5 3.5 3.5 2 4 4 4 5
– 15 20 25 20 22 25 25 20 25 30 25
504.9 490.7 395.5 398.6 562.5 478.1 390.4 433.5 614.2 441.8 391.7 563.2
± ± ± ± ± ± ± ± ± ± ± ±
Fig. 9. Experimentally measured longitudinal stresses in a laser reference weld (welding speed vw = 25 mm/s, beam power Pw = 5 kW) on the top and bottom surface compared to a specimen welded with the same parameters after heat treatment with the following parameters: vhf = 5 mm/s, beam power for heat treatment Phf = 4 kW, defocused beam radius r = 15.5 mm, dy = 30 mm.
reduction on the bottom surface of more than 70% is much higher than that on the top surface of less than 40%. The results given in Tables 2 and 3 for other radii show the same behaviour. It can be explained by two effects: the restored elastic energy (Eel = ·V) in the weld and the deformation of the welded and heat treated plates. The restored energy on the top surface is nearly twice the energy on the bottom surface because the beam weld on the top surface
49.7 33.4 73.0 27.0 82.3 60.5 75.9 30.1 58.1 49.2 26.7 34.7
Stress reduction on the top side [%]
Longitudinal stresses in the weld on the bottom side [MPa]
Stress reduction on the bottom side [%]
– 3 22 21 −1 5 23 14 −22 12 22 −12
482.0 ± 45.3 – – 339.2 ± 18.5 – – 202.0 ± 19.0 439.9 ± 25.5 – 209.5 ± 36.5 128.3 ± 39.6 18.2 ± 22.2
– – – 30 – – 58 9 – 57 73 96
Fig. 10. Cross section of an electron beam reference weld (welding speed vw = 15 mm/s, beam power Pw = 6.12 kW).
has double the width compared to the bottom surface as seen in the weld cross section represented in Fig. 10. The second reason for the larger stress reduction on the bottom surface is the deformation of the plate after welding and heat treatment. In Fig. 11 the distortion of a plate after welding (Fig. 11 left) and after welding and heat treatment (Fig. 11 right) are schematically shown. The welded plate shows almost no distortion. In case of an additional
Fig. 11. Scheme of the deformations of a plate after welding (left) and after welding and heat treatment (right), lengths in mm.
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process parameters. But the comparison of experimental and simulation results in Fig. 12 illustrates that this simplified simulation model is not suitable for the prediction of the real stress profiles. A more detailed model might provide a more precise prediction of the stress level. The effect of an enhanced heat penetration due to using lower travel speeds for the heat treatment can be used for thicker materials. Another possibility to increase the heat penetration is enabled by increasing the radius of the defocused beam. Both ways extend the exposure time of the heating process and result in smaller differences of the temperatures on the top and the bottom surface. 6. Conclusions
Fig. 12. Experimentally measured longitudinal stresses in a laser beam weld (welding speed vw = 25 mm/s, beam power Pw = 5 kW) treated with heating fields (vhf = 5 mm/s, beam power for heat treatment Phf = 4 kW, defocused beam radius r = 15.5 mm, dy = 30 mm) on the top and bottom surface compared to the corresponding two-dimensional simulation result.
heat treatment small transversal distortions of 1.4 mm and longitudinal distortions of 0.5 mm were measured with a profilometer. The longitudinal bulge is a reason for the larger stress reduction on the bottom surface. During the heating the top surface to temperatures of about 700 ◦ C the material on the top surface was deformed plastically due to the stress relaxation of the induced compressive stresses. On the bottom surface the induced compressive thermal stresses (in a temperature region of about 400 ◦ C) are too low for stress relaxing. So the plate bulges. During cooling the heat treated regions shrink and the bulging wanes in a small order but largely remains. This bending/bulging effect is an explanation of the larger stress reduction in the weld on the bottom surface compared to the top surface. Accordingly, the mechanical effect of additional tensile stress regions is much higher on the bottom surface. By applying larger heat source radii, the longitudinal stresses on the bottom side can be reduced further due to longer heat input resulting in higher heat penetration. Laser beam experiments with much lower temperature Tpost,max generated on the top surface than 700 ◦ C by using lower beam powers, for example Tpost,max ≈ 450 ◦ C at 2 kW beam power (instead of 4 kW for generating Tpost,max ≈ 700 ◦ C) and a beam radius of 15.5 mm, resulted in lower stress reductions (cf. Table 3). In this example the stress reduction on the top side is about 15% and on the bottom side only around 9%. The reason for the low stress reduction on the bottom side in this instance is the low heating of the bottom surface to about 200 ◦ C, which results in less plastic deformation generating less tensile stresses during cooling. For higher Tpost,max than 700 ◦ C on the top surface (cf. Table 3 beam radius of 15.5 mm and a beam power of 5 kW) the induced temperature difference between the top and the bottom surface is in a similar magnitude of about 300 ◦ C. So if the generated Tpost,max is higher than 700 ◦ C on the top surface, the temperature on the bottom surface is higher than 400 ◦ C, which results in more plastic deformation and a larger stress reduction during cooling on the bottom surface. In the experiment with 5 kW beam power and a beam radius of 15.5 mm in Table 3 the generated temperature Tpost,max of about 770 ◦ C on the top surface results in no stress reduction on the top surface of the weld but due to the higher generated temperatures on the bottom surface the stress reduction on the bottom side is increased to about 96%. The same tendency was found for the experiments with the electron beam in Table 2 with a beam radius of 6.5 mm, a travel speed of the heat treatment of 5 mm/s and different beam powers and for the simulation results. Hence, the adiabatic two-dimensional model is a useful tool for determining optimal
The presented method of mitigating welding residual stresses provides the potential to decrease longitudinal stresses in narrow welds with high stress gradients like in beam welds without any contact surfaces or additional equipment. Longitudinal stress reduction in this method is based on a mechanical unloading effect of the weld due to the additional tensile stress zones on both sides of the weld. The experimental results verify the assumptions made for the process mechanism and the process parameters based on the simulation results and substantiate that this process offers residual stress reductions of up to about 70% for 5 mm thick S355J2+N. Dependent on the component geometry and on the beam power it is possible to use different process parameters. The most suitable parameters are: generated temperatures slightly above the stress relief annealing temperature for ferritic steels of about 700 ◦ C, beam radii as large as possible (limited by the experimental setup), transversal distances of around the radius plus 11–15 mm. Acknowledgements The authors would like to thank the German Federation of Industrial Research Associations (AiF Arbeitsgemeinschaft industrieller Forschungsvereinigungen) and the German Federal Ministry for Trade, Industry and Technology (BMWi Bundesministerium für Wirtschaft und Technologie), for making this research possible by funding it in the Project 16139N “Anwendung der Mehrstrahltechnik zur Reduzierung der Eigenspannungen bei EB- und LB-geschweißten Bauteilen“(“Application of the multiplebeam-technique for residual stress reduction in EB and LB welded components”). In addition the authors would like to thank the HIGHYAG Lasertechnologie GmbH for providing their laser welding system for this research. References Brauss, M.E., Pineault, J.A., Eckersley, J.S., 1998. Residual stress characterization of welds and post-weld processes using X-ray diffraction techniques. In: Bossi, R.H., Pepper, D.M. (Eds.), Process Control and Sensors for Manufacturing (Proceedings of the SPIE – The International Society for Optical Engineering) 3399, San Antonio, USA. , pp. 196–204. ESI Group, 2006. Sysweld 2006 – Werkstoffkennwertdatenbank. Gebhardt, M.O., Quiroz, V., Gumenyuk, A., Rethmeier, M., 2010. Restraint effects on stresses and strains in single-run high power laser beam welding of thick plates. In: Cerjak, H., Enziger, N. (Eds.), Mathematical Modelling of Weld Phenomena 9. Verlag der Technischen Universität Graz, pp. 1011–1033. Guan, Q., Zhang, C.X., Guo, D.L., 1994. Dynamic control of welding distortion by moving spot heat sink. Welding in the World 33, 308–312. Guan, Q., Guo, D.L., Zhang, C.X., Li, J., 2006. Low stress no distortion welding based on thermal tensioning effects for thin materials. The Paton Welding Journal 12, 2–11. Han, S., Lee, T., Shin, B., 2002. Residual stress relaxation of welded steel components under cyclic load. Steel Research 73, 414–420. Mochizuki, M., 2007. Control of welding residual stress for ensuring integrity against fatigue and stress–corrosion cracking. Nuclear Engineering and Design 237, 107–123.
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Nitschke-Pagel, T., Dilger, K., 2006. Eigenspannungen in Schweißverbindungen – Teil 1: Ursachen der Eigenspannungsentstehung beim Schweißen. Schweißen und Schneiden 58, 466–479. Nitschke-Pagel, T., Dilger, K., 2007. Eigenspannungen in Schweißverbindungen – Teil 2: Bewertung von Eigenspannungen. Schweißen und Schneiden 59, pp. 23–32. Nitschke-Pagel, T., Dilger, K., 2007. Eigenspannungen in Schweißverbindungen – Teil 3: Verringerung von Eigenspannungen. Schweißen und Schneiden 59, pp. 387–395. Rosen, C.J., Gumenyuk, A., Zhao, H., Dilthey, U., 2007. Influence on residual stresses in electron beam welding. Science and Technology of Welding and Joining 12, 614–619.
Spieß, L., Teichert, G., Schwarzer, R., Behnken, H., Genzel, C., 2005. Röntgenographische Spannungsanalyse, In: Moderne Röntgenbeugung, B. G. Teubner Verlag 30, 5–382. Sun, Z., 1993. Residual stress in laser welded dissimilar steel tube-to-tube joints. Scripta Metallurgica et Materialia 29, 633–637. van der Aa, E.M., Hermans, M.J.M., Richardson, I.M., 2007. Control of welding residual stress and distortion by addition of a trailing heat sink. In: Cerjak, H., Bhadeshia, H.K.D.H., Kozeschnik, E. (Eds.), Mathematical Modelling of Weld Phenomena 8. Verlag der Technischen Universität Graz, pp. 1053–1072.
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Journal of Materials Processing Technology journal homepage: www.elsevier.com/locate/jmatprotec
Welding residual stress reduction by scanning of a defocused beam F. Tölle a,∗ , A. Gumenyuk a , A. Backhaus b , S. Olschok b , M. Rethmeier a , U. Reisgen b a b
BAM Federal Institute for Materials Research and Testing, Unter den Eichen 87, 12205 Berlin, Germany Welding and Joining Institute, RWTH Aachen University, Pontstrasse 49, 52062 Aachen, Germany
a r t i c l e
i n f o
Article history: Received 6 May 2011 Received in revised form 26 July 2011 Accepted 31 July 2011 Available online 6 August 2011 Keywords: Residual stresses Stress reduction High energy beam welding Post-weld heat treatment Laser scanner optics
a b s t r a c t The residual stresses in narrow electron or laser beam welds with high stress gradients are decreased without any contact surfaces or additional equipment by applying the welding beam after welding in a defocused mode for heating the material regions in a certain distance from the weld on both sides. In case of electron beam application, the beam is positioned and focused by the electromagnetic coil with high frequency. In case of laser beam application a laser scanner optics enables fast positioning by an optomechanic beam deflection, while defocusing of the laser beam is obtained by increasing the distance between scanner optics and workpiece. Dependent on the component geometry and on the beam power different process parameters are used. The adjustable process parameters are the radius and the power of the defocused beam and the transversal and longitudinal distances between the welding and the defocused beam. The mechanism and the influence of the process parameters are investigated by FEM-simulation and a number of experiments on a ferritic steel S355J2+N with 5 mm thickness. FEMsimulation is used to reduce the matrix of process parameters for the experiments. The best experimental result shows a stress reduction of about 70%. © 2011 Elsevier B.V. All rights reserved.
1. Introduction In welding, a small material region is melted and cooled down with high temperature gradients. As described by Nitschke-Pagel and Dilger (2006) in their review of the causes of welding residual stresses for diverse welding procedures and various materials, the heat-affected material shrinks during cooling and causes stresses. Especially due to high energy welding processes like electron or laser beam welding these stresses can reach the temperature and phase dependent local yield strength in longitudinal weld direction. Stresses ranging at this level will be relaxed by plastic deformations as shown by Nitschke-Pagel and Dilger (2007a) in their review concerning welding residual stress assessment. The authors point out that residual stresses can influence the performance especially for high-strength materials. In consequence of stress relaxation or stress corrosion cracking these high stresses can limit service life depending on the tensile strength of materials as discussed by Nitschke-Pagel and Dilger (2007a) referring to a number of fatigue tests. Han et al. (2002) substantiated this influence of welding residual stresses on the fatigue strength as well as on the component life and developed a model for quantitative prediction of residual stress relaxation under cyclic load. The influence of high tensile residual
∗ Corresponding author. Tel.: +49 30 8104 3101; fax: +49 30 8104 1557. E-mail address: fl[email protected] (F. Tölle). 0924-0136/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.jmatprotec.2011.07.019
stresses on stress corrosion cracking was discussed by Sun (1993) and Brauss et al. (1998), too. A large number of processes for lowering the welding residual stresses is available like stress relief annealing as mentioned and described by Nitschke-Pagel and Dilger (2007b) or the static low-stress-no-distortion-technique as developed and discussed by Guan et al. (2006) which works through generating a specified temperature profile in the welded component using heating and cooling elements. But these methods are better applicable for wider welds such as TIG welds and for simple component geometries to provide the required contact surfaces for heating and cooling elements. They are also cost-intensive. The more flexible dynamically controlled low-stress-no-distortion-technique invented and described by Guan et al. (1994) uses a trailing cooling jet for generating the required temperature profile. Nevertheless, van der Aa et al. (2007) proved that this method is only effective for cooling jet travel speeds of less than 8 mm/s. For faster welding processes like beam welding with ordinary welding speeds higher than 10 mm/s the penetration effect of this cooling is minimized. Other mechanical methods of stress reduction yield lower tensile stresses only in near-surface areas. Water jet peening is one of such methods. Mochizuki (2007) showed that this peening effect only introduces compressive stresses in the treated surface and enhances the fatigue strength due to the reduction of stress corrosion cracking. Beam welding processes offer the advantages of low component deformations, low heat input and remote application, which enable joining of complex component geometries. The previously
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Fig. 1. Process scheme for simulation and experimental investigation of a welding beam and heating fields produced with the important geometric process parameters: radius of the heat source r, longitudinal distance between welding source and heating fields dx and transversal distance between welding source and heating fields dy .
designed stress reducing techniques are not practicable for this fast remote application because of the required contact surfaces. Therefore, a research concept for a stress mitigating process applicable to beam welded components is presented in this paper aimed at developing a contactless technique operating without any additional equipment. As shown by Rosen et al. (2007), application of the electron or the laser beam after welding for the reduction of high longitudinal residual stresses enables such a remote process. In this method the beam is used in a defocused mode for heating the material regions on both sides of the weld from a certain distance to generate additional tensile stress zones which unload the weld. Due to this unloading effect the residual stresses in the weld are decreased. The process scheme of welding and heat treatment in one work stage, i.e. simultaneous welding and heat treatment, is shown in Fig. 1. It is also possible to perform the heat treatment in a second work stage separated from the welding process. The important investigated process parameters, which will be explained in this article, are the defocused beam radius r, the maximum temperature Tpost,max to be generated by the defocused beam in the heated regions, and the transversal dy and longitudinal dx distances between the weld centre and the defocused beam. In Rosen’s (2007) work only the feasibility of this process for stress reduction was investigated without an analysis and any explanation of the stress reduction mechanism in this method. Furthermore, these researches are restricted for material thicknesses of 2 mm and 3 mm, a defocused beam radius of 10 mm (achieved stress reduction of about 85%) and only for the electron beam application. At this point the investigations in this article continue Rosen’s work. In these researches the stress reduction mechanism is described and the whole process parameters were varied and investigated. In addition, this method was applied to a laser beam welding source. The process parameters correlate with each other and generate a huge parameter matrix for the experimental investigations. To reduce the dimension of this matrix for the experiments with an electron and a laser beam, respectively, a qualitative parameter study was performed by FEM-simulation. With the simulative and the experimental results advices for the most suitable parameter sets can be concluded. 2. Finite element simulation to investigate the influence of the process parameters For the parameter study with the ANSYS software a simplified simulation model had to be created. This model was only a tool for a qualitative investigation of the stress reduction mechanism and to define process parameter regions for the experimental researches. Based on the resulting symmetry conditions (cf. Fig. 1)
Fig. 2. Scheme of the dependence of the power density distribution of an electron beam from the position of the beam focus to the workpiece.
in the finite element model, only half of the specimen was simulated to decrease simulation time. Another simplification was the use of a two-dimensional concept for the thermal and the structural model with defined z-depth characterizing the specimen thickness of 5 mm. Thus, the adiabatic thermal model represents ideal heat conduction in the heating fields caused by the defocused beam in thickness direction (z-direction). The adiabatic model was used because the two-dimensional model should represent a layer in the middle of the plates. In this model no heat transfer in vertical direction takes place and this corresponds to the adiabatic boundary conditions on the upper and the lower surfaces. The cooling of this layer due to heat conduction in thickness direction and due to the exchange with the environment is complex. So for this qualitative investigation, which parameter regions are most suitable for the stress reduction, it is not important to cool down the simulated plates to room temperature. A further cooling would result in an increase of the stresses but do not influence the most qualified parameter regions for the experimental investigations because the stress reducing mechanism is induced during the heat treatment and not during cooling down after this heat treatment. In the structural model the plane stress condition is illustrated by the two-dimensional layout. The elastoplastic material model was designed based on the temperature dependent material properties of S355J2+N from the SYSWELD database (ESI Group, 2006). To consider phase transformation during the welding process, a peak temperature driven approach as explained by Gebhardt et al. (2010) was used. This approach based on a change of the material properties if the element temperatures reach specific values. If they are larger than 740 ◦ C the martensitic material model for the yield strengths and other thermal expansions were used. The heat source for welding was modelled as a traveling point heat source and for the post welding heat treatment representing the defocused beam a cyclic area with a specified radius and a homogeneously distributed power density was modelled. This approach for the power density distribution correlates with measurements of a defocused electron beam in over focus position, as seen in Fig. 2. In Fig. 2 the dependence of the power density distribution of an electron beam from the position of the beam focus to the workpiece is shown. With the distribution in over focus position the smoothest temperature gradient in the heat treated areas is generated. So this over focus position was used for the experimental investigations with the electron beam welding source and for modelling the defocused heat source.
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Fig. 3. Simulated longitudinal stress distribution in the workpiece across to the welding direction after welding (solid line) and after applied heat treatment (dashed lines) with different temperatures Tpost,max generated by the defocused beam influencing the stress reduction in the weld with the process parameters: r = 15 mm, dy = 25 mm, dx = 200 mm and vhf = 15 mm/s.
This homogeneously distribution of the power density does not represent the distribution in a defocused laser beam. Therefore, the conclusions for the most suitable process parameters resulting from the simulation and the experiments with the electron beam were adapted to the experiments with the laser beam and checked for their applicability. With this FEM-model it is possible to investigate the qualitative influence of the process parameters. 3. Results of the finite element simulations Fig. 3 shows the influence of the temperature Tpost,max generated by the defocused beam on the stress reduction potential in the weld. This figure as well as Figs. 4 and 5 correspond to the heat treatment conducted in a second work stage (dx = weld length = 200 mm) separated from the welding process. During cooling, stresses in the range of around 600 MPa were generated in the weld (welding speed vw = 15 mm/s, beam power Pw = 800 W). The Tpost,max value was determined by the beam radius r, the travel speed vhf and the beam power Phf used for this heating field. Up to temperature of about 700 ◦ C, the stress reduction increases with rising temperatures. This behaviour can be explained by analysing the twodimensional stress distribution illustrated in Fig. 4 (with process parameters r = 10 mm, dy = 27 mm, dx = 200 mm, travel speed v = 5 mm/s and Tpost,max = 801 ◦ C) and the time dependent tem-
Fig. 4. Simulation result for the longitudinal stresses in MPa after heat treatment with the process parameters: r = 10 mm, dy = 27 mm, dx = 200 mm, vhf = 5 mm/s and Tpost,max = 801 ◦ C.
Fig. 5. Temperature (top) and longitudinal stress (bottom) profile over time in the weld (Tweld , weld ), in the centre of the heating field (THF , HF ) and in not heat treated region (TNHT , NHT ) with the process parameters: r = 10 mm, dy = 27 mm, dx = 200 mm, vhf = 5 mm/s and Tpost,max = 801 ◦ C (positions in the model are shown in Fig. 4; tbHT indicates the point in time of heat treatment, ttu indicates the point in time of termination of the unloading effect).
perature and stress profiles reflected in Fig. 5. In Fig. 4, the two-dimensional longitudinal stress profile is plotted after the heat treatment and a subsequent cooling time of 8000 s. In Fig. 5, the corresponding temperature profiles in the weld Tweld (cf. Fig. 4 coordinates x = 100 mm and y = 0 mm), in the heating fields THF (cf. Fig. 4 coordinates x = 100 mm and y = 27 mm) and in the non-influenced area TNHT (cf. Fig. 4 coordinates x = 100 mm and y = 84 mm) are correlated with the longitudinal stress profiles ( weld , HF , NHT ) in the same points. They were recorded over the time during processing on the simulated welded plate. After welding, high tensile longitudinal stresses weld were present in the weld. Due to heating by the defocused beam, the temperature dependent yield strength decreases in the heated region, and thus the generated compressive stresses (cf. Fig. 5 HF between 30 s and 35 s) can relax based on plastic deformations. The yield strength of low alloyed ferritic steels decreases with a high gradient in the temperature range between 600 ◦ C and 800 ◦ C during heating. If the generated temperature ranges below 700 ◦ C, the induced compressive stresses are lower and the yield strength in lower temperature regions is too high for stress relaxation. Temperatures above 800 ◦ C or 900 ◦ C give only marginal further yield strength reductions, i.e. the most suitable Tpost,max is about 700 ◦ C for the ferritic steel S355J2+N. While the heat treated regions cool down, the compressive stresses decrease and switch to tensile stresses due to material shrinkage. These additional tensile stress zones on both sides of the weld lead to a mechanical unloading of the weld as shown in Fig. 5 by weld in the time interval between 30 s and 70 s. The magnitude of the evolution of longitudinal stresses in the weld and of the unloading effect in dependence on the longitudinal distance dx between melting and post welding heat treatment processes can be seen in Fig. 6, too. This diagram shows the stress generation in the weld during cooling down from the melting temperature for an as welded plate compared to welded plates with a subsequent heat treatment at different dx while processing. The larger dx , the higher are the tensile stresses which can be generated in the cooling time interval (cf. Figs. 5 and 6 till the point in time before heat treatment tbHT ) between the welding and the reheating process. During cooling of the additionally heated regions down to
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Fig. 6. Evolution of longitudinal stresses in the weld dependent on the longitudinal distance between the welding source and the heating fields dx while cooling from melting temperature (tbHT indicates the point in time of heat treatment, ttu indicates the point in time of termination of the unloading effect); process parameters: r = 15 mm, dy = 25 mm, vhf = 15 mm/s and Tpost,max = 830 ◦ C.
the same temperature as that in the weld zone (cf. Figs. 5 and 6 for Tweld ≈ THF at the point in time of termination of unloading ttu = 60 s), the tensile stresses in the weld are decreased by unloading. In other words, the unloading occurs as long as the heat treated regions shrink more than the weld zone. As a result of heat conduction the weld is reheated. At the point in time ttu , the temperatures in the heated zones and in the weld are equal and the unloading terminates. It can also be seen in Fig. 6 that a shift to larger dx , i.e. longer cooling times before heat treatment tbHT , increases the magnitude of the unloading effect. Further cooling of the weld and the heat treated regions causes an increase in the longitudinal stresses. Due to the idealized conditions in the adiabatic two-dimensional model the workpieces exhibit a homogeneously distributed temperature of about 270 ◦ C at the end of the simulations after around 8000 s cooling time. For small dx (nearby 0 mm), the temperature in the weld is higher than in the additionally heat treated regions. In this case no mechanical unloading occurs through the heat treatment. If dx is so large that the weld is able to cool down to ambient temperature, a further increase of dx does not lead to significantly larger stress reductions. In further simulations, the influences of the transversal distance dy and of the defocused beam radius r on stress reduction were investigated. In this respect, the most suitable dy depending on the defocused beam radius was determined to obtain the maximal value of the stress reduction in the weld zone. Fig. 7 represents the maximum stresses in the weld depending on dy at different radii r. Each single point shows the value of the calculated longitudinal residual stress in the weld corresponding to a specific beam radius r and a certain transversal distance dy . From the diagram given in Fig. 7 it can be concluded that dy ≈ r + 12 mm provides the most suitable stress reduction of this process. With these simulation results the stress reduction mechanism can be illustrated and parameter regions which are most suitable for the experimental investigations can be defined. 4. Experimental For qualifying this process for stress reduction in the weld, a number of experiments with varied process parameters were carried out on 5 mm thick ferritic steel S355J2+N. Parameters resulting from the presented FEM-simulations were used. The applied workpieces were free from residual stresses before welding.
Fig. 7. Correlation of the radius of the heat source r with the transversal distance between welding source and heating fields dy (dx = 200 mm, vhf = 15 mm/s and Tpost,max ≈ 700 ◦ C), the longitudinal residual stresses in the weld centre (each point represents one simulation result) are shown (stress in the weld without additional heat treatment in grey at 580 MPa).
In the electron beam welding experiments, full penetration welds were produced by the focused beam with 0.3 mm beam radius and 6.12 kW beam power Pw . The welding speed vw was 15 mm/s for all experiments. Workpiece dimensions of 200 mm × 200 mm × 5 mm were used. Due to the feasibility of high frequency electromagnetic beam focusing and deflection in electron beam welding it is possible to use the beam for the heat treatment in the defocused mode simultaneously with welding. Therefore, the electron beam was positioned and deflected with a frequency of 500 Hz to the three positions in the following sequence: weld position, heat treatment position to the left of the weld, weld position, heat treatment position to the right of the weld, back to the weld position. For welding and heat treatment in one work stage the same speed for all three positions must be used (vw = vhf = 15 mm/s). Due to the machine dependent maximum deflection distance of 70 mm, experiments with different travel speeds for welding and heat treatment (vw = / vhf ) as well as experiments with longitudinal distances larger than 70 mm had to be performed in a second work stage. In the cases of two work stages, the focused beam was used in the first work stage for the welding process with the same welding parameters as applied to the in situ heat treatment process. In the second work stage the defocused beam was deflected between the two positions of heat treatment with a frequency of 500 Hz, too, and moved along the weld with a travel speed vhf of 15 mm/s. In these cases, the weld was allowed to cool down to temperatures below 200 ◦ C before the heating process was started in a second work stage. Under these conditions, the time variable tx describing the delay between the end of welding and the beginning of heat treatment was used instead of the longitudinal distance dx . The different beam radii for the heat treatment were generated by varying the focus current and controlled by measurements of the power density distribution. For three radii (6.5 mm, 10.0 mm and 15.0 mm) aligned powers Phf (2.88 kW, 3.12 kW and 5.28 kW) of the defocused beam were used for heating the workpiece surface. The parameter regions investigated in the electron beam experiments are listed in Table 1. The second part of the experimental program was carried out with laser beam welding using a scanner optics that operates by optomechanic beam focusing and deflection. However, as the deflection speed of the laser optics was limited to 14 Hz and was much slower than the electromagnetic deflection of the electron beam, it was not possible to conduct keyhole welding with simultaneous heat treatment at different positions in one work stage. In
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Table 1 Investigated parameter regions in the electron beam experiments. Radius of the defocused beam [mm]
dy [mm]
tx [s]
Travel speed vhf [mm/s]
Power for defocused beam Phf [kW]
6.5 10.0 15.0
0, 5, 10, 15, 20, 30 23 26
0–4.66, 18–43 18, 30 18, 30
2.5, 5, 15 5, 15 5, 15
0.48–2.52 1.25, 3.12 2.11, 5.28
this case, heat treatment in laser beam processes had to be performed in a second work stage. For these experiments a 5 kW fibre laser coupled with the laser scanner optics was used with the following welding parameters: 5 kW beam power Pw , 0.3 mm beam radius of the focused laser beam and 25 mm/s welding speed vw . To obtain different radii of the defocused beam, the beam waist was positioned in specific distances above the specimen surface. The three different radii 9.5 mm, 12.0 mm and 15.5 mm (including 86% of the beam power of the defocused beam) in the workpiece plane were controlled by measurements with a power density meter. After welding and cooling down to ambient temperature, the second work stage was executed by lifting the scanner optics to the adjusted distance and using different powers for the specific beam radius at a travel speed vhf of 5 mm/s. Accordingly, for laser beam experiments it was only possible to investigate the effect of this process when the temperature in the weld was low. This means that the variable tx adopted to approach infinity and was not specified for the individual experiments. The duration of 60 ms was used for heating one position before the beam was deflected to the other side of the weld for heating. With the help of a thermographic camera the necessary power values Phf for heating the workpiece surface in the heating field centre to about 700 ◦ C were aligned to obtain corresponding values for the defocused beam radius. Using X-ray diffraction, the residual stresses were measured in the workpiece surfaces (averaged over the X-ray penetration depth of up to 100 m) after the experiments. Calculation of the stresses was carried out adopting the sin2 -method as described by Spieß et al. (2005).
Fig. 8. Experimentally measured longitudinal stresses in an electron beam reference weld (welding speed vw = 15 mm/s, beam power Pw = 6.12 kW) on the top and bottom surface compared to a specimen welded with the same parameters after heat treatment with the following parameters: vhf = 15 mm/s, beam power for heat treatment Phf = 5.28 kW, defocused beam radius r = 15 mm, dy = 26 mm, time interval between welding source and heating fields tx = 30 s.
higher laser power levels and travel speeds for heating the material surface to this temperature region are not possible. Table 3 lists the best results of the X-ray diffraction measurements in the laser beam investigations showing similar stress reductions compared to the electron beam experiments. The longitudinal stress profile of the experiment with the best stress reducing result of 22% on the top and 73% on the bottom surface and the radius of 15.5 mm is compared to the reference weld in Fig. 9. The results listed in Tables 2 and 3 verify the simulation results concerning the influence and the values of the most suitable process parameters. In the experiments with both beam sources similar results were generated for the most suitable parameters. Accordingly, these results were compared together with the simulation results. As assumed from the FEM-simulations, in Tables 2 and 3 the lowest experimentally measured residual stresses in the weld for each radius are achieved with a dy value of about the radius plus 11–15 mm (hypothesis from simulation results dy ≈ r + 12 mm). In Figs. 8 and 9 the longitudinal stress profiles for the best experienced parameter sets are compared with reference welds for electron and laser beam welds. Remarkably, for both beam sources the stress
5. Experimental results and discussion of the stress reducing capability The best results obtained in the electron beam investigations are listed in Table 2. The results of the X-ray diffraction measurements reveal for the largest experienced radius of 15 mm a stress reduction of about 36% on the weld top surface and of 76% on its bottom surface. The longitudinal stress profile of this specimen with the best stress mitigating result is compared to the stress profile in a reference weld in Fig. 8. In the experiments with the laser beam, the desired temperatures in the heating fields of around 700 ◦ C for the defocused laser beam radius of 15.5 mm were reached with 4 kW beam power and 5 mm/s travel speed. Due to the maximum scanner capacity (5 kW)
Table 2 Longitudinal stresses in the electron beam weld centre reference weld without heat treatment. Radius of the defocused beam [mm]
Beam power for heat treatment Phf [kW]
vhf [mm/s]
dy [mm]
tx [s]
Longitudinal stresses in the weld on the top side [MPa]
Reference 6.5 6.5 6.5 6.5 6.5 6.5 10.0 15.0
– 2.88 2.88 2.88 0.96 1.12 1.44 3.12 5.28
– 15 15 15 5 5 5 15 15
– 10 20 35 20 20 20 23 26
– 5 5 5 5 5 5 30 30
471.4 430.3 384.7 500.0 468.3 374.8 440.2 377.0 304.1
± ± ± ± ± ± ± ± ±
13.0 32.3 60.6 54.1 63.3 45.1 61.9 58.3 53.8
Stress reduction on the top side [%]
Longitudinal stresses in the weld on the bottom side [MPa]
Stress reduction on the bottom side [%]
– 9 18 −6 1 20 7 20 36
375.1 ± 12.7 – 354.8 ± 23.0
– – 5
– – – 89.3 ±59.0 91.9 ±43.6
– – – 76 76
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Table 3 Longitudinal stresses in the middle of the laser beam weld (vhf = 5 mm/s). Radius of the defocused beam [mm]
Beam power for heat treatment Phf [kW]
dy [mm]
Longitudinal stresses in the weld on the top side [MPa]
As welded 9.5 9.5 9.5 12.0 12.0 12.0 15.5 15.5 15.5 15.5 15.5
– 2.3 2.3 2.3 3.5 3.5 3.5 2 4 4 4 5
– 15 20 25 20 22 25 25 20 25 30 25
504.9 490.7 395.5 398.6 562.5 478.1 390.4 433.5 614.2 441.8 391.7 563.2
± ± ± ± ± ± ± ± ± ± ± ±
Fig. 9. Experimentally measured longitudinal stresses in a laser reference weld (welding speed vw = 25 mm/s, beam power Pw = 5 kW) on the top and bottom surface compared to a specimen welded with the same parameters after heat treatment with the following parameters: vhf = 5 mm/s, beam power for heat treatment Phf = 4 kW, defocused beam radius r = 15.5 mm, dy = 30 mm.
reduction on the bottom surface of more than 70% is much higher than that on the top surface of less than 40%. The results given in Tables 2 and 3 for other radii show the same behaviour. It can be explained by two effects: the restored elastic energy (Eel = ·V) in the weld and the deformation of the welded and heat treated plates. The restored energy on the top surface is nearly twice the energy on the bottom surface because the beam weld on the top surface
49.7 33.4 73.0 27.0 82.3 60.5 75.9 30.1 58.1 49.2 26.7 34.7
Stress reduction on the top side [%]
Longitudinal stresses in the weld on the bottom side [MPa]
Stress reduction on the bottom side [%]
– 3 22 21 −1 5 23 14 −22 12 22 −12
482.0 ± 45.3 – – 339.2 ± 18.5 – – 202.0 ± 19.0 439.9 ± 25.5 – 209.5 ± 36.5 128.3 ± 39.6 18.2 ± 22.2
– – – 30 – – 58 9 – 57 73 96
Fig. 10. Cross section of an electron beam reference weld (welding speed vw = 15 mm/s, beam power Pw = 6.12 kW).
has double the width compared to the bottom surface as seen in the weld cross section represented in Fig. 10. The second reason for the larger stress reduction on the bottom surface is the deformation of the plate after welding and heat treatment. In Fig. 11 the distortion of a plate after welding (Fig. 11 left) and after welding and heat treatment (Fig. 11 right) are schematically shown. The welded plate shows almost no distortion. In case of an additional
Fig. 11. Scheme of the deformations of a plate after welding (left) and after welding and heat treatment (right), lengths in mm.
F. Tölle et al. / Journal of Materials Processing Technology 212 (2012) 19–26
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process parameters. But the comparison of experimental and simulation results in Fig. 12 illustrates that this simplified simulation model is not suitable for the prediction of the real stress profiles. A more detailed model might provide a more precise prediction of the stress level. The effect of an enhanced heat penetration due to using lower travel speeds for the heat treatment can be used for thicker materials. Another possibility to increase the heat penetration is enabled by increasing the radius of the defocused beam. Both ways extend the exposure time of the heating process and result in smaller differences of the temperatures on the top and the bottom surface. 6. Conclusions
Fig. 12. Experimentally measured longitudinal stresses in a laser beam weld (welding speed vw = 25 mm/s, beam power Pw = 5 kW) treated with heating fields (vhf = 5 mm/s, beam power for heat treatment Phf = 4 kW, defocused beam radius r = 15.5 mm, dy = 30 mm) on the top and bottom surface compared to the corresponding two-dimensional simulation result.
heat treatment small transversal distortions of 1.4 mm and longitudinal distortions of 0.5 mm were measured with a profilometer. The longitudinal bulge is a reason for the larger stress reduction on the bottom surface. During the heating the top surface to temperatures of about 700 ◦ C the material on the top surface was deformed plastically due to the stress relaxation of the induced compressive stresses. On the bottom surface the induced compressive thermal stresses (in a temperature region of about 400 ◦ C) are too low for stress relaxing. So the plate bulges. During cooling the heat treated regions shrink and the bulging wanes in a small order but largely remains. This bending/bulging effect is an explanation of the larger stress reduction in the weld on the bottom surface compared to the top surface. Accordingly, the mechanical effect of additional tensile stress regions is much higher on the bottom surface. By applying larger heat source radii, the longitudinal stresses on the bottom side can be reduced further due to longer heat input resulting in higher heat penetration. Laser beam experiments with much lower temperature Tpost,max generated on the top surface than 700 ◦ C by using lower beam powers, for example Tpost,max ≈ 450 ◦ C at 2 kW beam power (instead of 4 kW for generating Tpost,max ≈ 700 ◦ C) and a beam radius of 15.5 mm, resulted in lower stress reductions (cf. Table 3). In this example the stress reduction on the top side is about 15% and on the bottom side only around 9%. The reason for the low stress reduction on the bottom side in this instance is the low heating of the bottom surface to about 200 ◦ C, which results in less plastic deformation generating less tensile stresses during cooling. For higher Tpost,max than 700 ◦ C on the top surface (cf. Table 3 beam radius of 15.5 mm and a beam power of 5 kW) the induced temperature difference between the top and the bottom surface is in a similar magnitude of about 300 ◦ C. So if the generated Tpost,max is higher than 700 ◦ C on the top surface, the temperature on the bottom surface is higher than 400 ◦ C, which results in more plastic deformation and a larger stress reduction during cooling on the bottom surface. In the experiment with 5 kW beam power and a beam radius of 15.5 mm in Table 3 the generated temperature Tpost,max of about 770 ◦ C on the top surface results in no stress reduction on the top surface of the weld but due to the higher generated temperatures on the bottom surface the stress reduction on the bottom side is increased to about 96%. The same tendency was found for the experiments with the electron beam in Table 2 with a beam radius of 6.5 mm, a travel speed of the heat treatment of 5 mm/s and different beam powers and for the simulation results. Hence, the adiabatic two-dimensional model is a useful tool for determining optimal
The presented method of mitigating welding residual stresses provides the potential to decrease longitudinal stresses in narrow welds with high stress gradients like in beam welds without any contact surfaces or additional equipment. Longitudinal stress reduction in this method is based on a mechanical unloading effect of the weld due to the additional tensile stress zones on both sides of the weld. The experimental results verify the assumptions made for the process mechanism and the process parameters based on the simulation results and substantiate that this process offers residual stress reductions of up to about 70% for 5 mm thick S355J2+N. Dependent on the component geometry and on the beam power it is possible to use different process parameters. The most suitable parameters are: generated temperatures slightly above the stress relief annealing temperature for ferritic steels of about 700 ◦ C, beam radii as large as possible (limited by the experimental setup), transversal distances of around the radius plus 11–15 mm. Acknowledgements The authors would like to thank the German Federation of Industrial Research Associations (AiF Arbeitsgemeinschaft industrieller Forschungsvereinigungen) and the German Federal Ministry for Trade, Industry and Technology (BMWi Bundesministerium für Wirtschaft und Technologie), for making this research possible by funding it in the Project 16139N “Anwendung der Mehrstrahltechnik zur Reduzierung der Eigenspannungen bei EB- und LB-geschweißten Bauteilen“(“Application of the multiplebeam-technique for residual stress reduction in EB and LB welded components”). In addition the authors would like to thank the HIGHYAG Lasertechnologie GmbH for providing their laser welding system for this research. References Brauss, M.E., Pineault, J.A., Eckersley, J.S., 1998. Residual stress characterization of welds and post-weld processes using X-ray diffraction techniques. In: Bossi, R.H., Pepper, D.M. (Eds.), Process Control and Sensors for Manufacturing (Proceedings of the SPIE – The International Society for Optical Engineering) 3399, San Antonio, USA. , pp. 196–204. ESI Group, 2006. Sysweld 2006 – Werkstoffkennwertdatenbank. Gebhardt, M.O., Quiroz, V., Gumenyuk, A., Rethmeier, M., 2010. Restraint effects on stresses and strains in single-run high power laser beam welding of thick plates. In: Cerjak, H., Enziger, N. (Eds.), Mathematical Modelling of Weld Phenomena 9. Verlag der Technischen Universität Graz, pp. 1011–1033. Guan, Q., Zhang, C.X., Guo, D.L., 1994. Dynamic control of welding distortion by moving spot heat sink. Welding in the World 33, 308–312. Guan, Q., Guo, D.L., Zhang, C.X., Li, J., 2006. Low stress no distortion welding based on thermal tensioning effects for thin materials. The Paton Welding Journal 12, 2–11. Han, S., Lee, T., Shin, B., 2002. Residual stress relaxation of welded steel components under cyclic load. Steel Research 73, 414–420. Mochizuki, M., 2007. Control of welding residual stress for ensuring integrity against fatigue and stress–corrosion cracking. Nuclear Engineering and Design 237, 107–123.
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