RSW of Dissimilar Steel
This article is the summary of a paper entitled, “Weld Nugget Penetration of a Four-Sheet Resistance Spot Welding Advanced High-Strength Steels”, by K. Namola, et al.N-11
Experimental Weld Nugget Penetration
The study analyzes the effect of electrode size and composition on final weld nugget size and penetration. Nugget growth patterns were analyzed and weldability issues characterized. Figure 1 shows the arrangement of the four-layer stack-ups that were tested in this study. Truncated code electrodes used were a 6-mm Class 1, 6-mm Class 3, 6-mm Class 20, 8-mm Class 1, and 10-mm Class 1. Samples were welded in the as-received condition. JAC270 is a cold rolled Mild steel with a galvanneal coating having a minimum tensile strength of 270 MPa. JSC590 and JSC980 are bare cold rolled Dual Phase steels with a minimum tensile strength of 590 MPa and 980 MPa, respectively.
Figure 1: Resistance Welding Stack and Test Electrode Combinations.N-11
Best results from the iterative trials were obtained using an 8- and 6-mm Class 1 copper electrode with the weld schedule shown in Figure 2. This weld schedule was repeated using the electrode combinations listed in Table 1. Figure 3 shows cross sections of each weld listed in Table 1.
Figure 2: Down-Selected Weld Schedule from Trials.N-11
Table 1: Nugget Penetration Using the Down-Selected Weld Schedule from Trials and Different.N-11
Figure 3: Welds Made Using the Down-Selected Schedule and Different Electrodes.N-11
Figure 4 shows cross sections of five welds made starting with new 8- and 6-mm Class 1 electrodes. As can be seen, expulsion gets progressively worse over time but penetration does not. Penetration values into the JAC 270 were determined by metallography and are shown in Table 2.
Figure 4: Welds from Repeatability Study Using the Down-Selected Weld Schedule and 8- and 6-mm Class 1 Electrodes.N-11
Table 2: Nugget Penetration into the JAC 270 During Repeatability Study.N-11
Table 3 lists the resistance measurements at the weld stack interfaces. Figure 5 shows the resistance graph of weld stack up.
Table 3: Resistance Measurements of Weld Materials and Weld Stack Interfaces.N-11
Figure 5: Resistance Graph of Weld Stack-Up.N-11
The weld force used was 2.3 kN and the current reduction values are listed in Table 4. Figure 6 and Figure 7 show the still images at each pulse. The heating pattern implies that the JAC270 is forged into the weld nugget.
Table 4: Current Reduction for High-Speed Video Welds.N-11
Figure 6: Still Images from High-Speed Video.N-11
Figure 7: Still Images from Weld Simulation of the Down-Selected Schedule using 8- and 6-mm Class 1 Electrodes.L-58
Press Hardened Steels, Solid State Welding
This article summarizes a paper, entitled “Effect of GA-Coating Evolution during Press-Hardening on Fiber Laser Lap Welding Behavior of 22MnB5 Steel”, by M. H. Razmpoosh, et al.R-4
The study investigates the effects of Fe-Zn diffusion layer on laser lap-joining behavior of galvanneal (GA) coated 22MnB5 steel, an Advanced High-Strength Steel designed for the hot forming process. The results indicate that by using higher press-hardening durations, the weld window shrinks; however, this results in a wider weld bead, and therefore promotes the load-bearing capacity of the joint.
Press-hardened 22MnB5, 2mm sheet steels were used in the present study. The details of the chemical composition and the as-received mechanical properties of the sheets are given in Table 1. The steel sheets were GA-coated with two different initial total coating weights of 100 and 140 g/m2 (Table 2).
Table 1: Chemical Composition (wt.%) and Mechanical Properties of the Experimental PHS.
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Table 2: Weight and Chemical Composition of Various GA Coatings used in the Present Study.
Figure 1 demonstrates backscattered scanning electron microscopy (BS-SEM) and Electron probe microanalysis (EPMA) elemental distribution of a representative Fe-Zn DL after press-hardening at 860°C for 4-10 min and corresponding 900°C for 10 min. It has been observed that by increasing the press-hardening time the Zn-content decreases; however, at higher press-hardening temperatures (i.e., 900°C) due to extreme oxidation, the average Zn-content decreases severely.
![Figure 1. BS-SEM and EPMA Results of the Press-Hardened Blanks at 860°C [(a) 4 min, (b) 7 min, (c) 10 min, and (d) 900°C for 10 min (DL)].](https://ahssinsights.org/wp-content/uploads/2020/07/3112_Fig1.jpg)
Figure 1. BS-SEM and EPMA Results of the Press-Hardened Blanks at 860°C [(a) 4 min, (b) 7 min, (c) 10 min, and (d) 900°C for 10 min (DL)].
Figure 2 summarizes the effects of press-hardening time and temperature on the thickness of the Fe-Zn DL in two different initial coating weights of 100 and 140 g/m2. With increasing the heat-treatment time at 860°C, the thickness of overall Fe-Zn DL increases. However, specifically at 900°C and longer press-hardening times, the final Fe-Zn DL is not increasing. Moreover, it has been observed that at a constant press-hardening time-temperature, lower initial coating weight results in a lower final Fe-Zn DL thickness.
Figure 2: DL Thickness vs. Press-Hardening Times at the Experimental Temperature and Initial Coating Weights.
According to Figure 3(a), increasing the press-hardening time at a constant temperature results in wider weld beads. Hence, the fact that failure occurs within the FZ (faying surface) during lap-shear tensile tests justifies the slightly enhanced peak loads [Figure 3(b)].
Figure 3: (a) Joint Width, (b) Peak Load of Lap-Shear Tensile Test vs. Press-Hardening Time, and (c) Schematic of Fe-Zn DL and Laser Interaction.
This work concluded the following:
- The initial GA-coating mainly evolves into a Fe-Zn DL [α-Fe(Zn)] and ZnO after the press-hardening. The thick α-Fe(Zn) phase holding 20-40% Zn; however, it was observed that with increasing press-hardening temperature, due to severity of oxidation Zn-content of the Fe-Zn DL decreases.
- Due to higher oxidation, severity at higher press-hardening temperatures, and subsequent lower Zn-content, the sensitivity of the process window is less than 860°C.
- Because of intensified Zn-plasma and laser beam interaction, by increasing the press hardening time at a constant temperature of 860°C (higher Fe-Zn DL thickness), joint width increases. This explains higher lap-shear tensile peak loads associated with the higher press-hardening times.
RSW of Dissimilar Steel
This article summarizes a paper entitled, “Higher than Expected Strengths from Dissimilar Configuration Advanced High-Strength Steel Spot Welds”, by E. Biro, et al.B-6
This study shows that the cross tension strength (CTS) is always higher than the strength expected from the lower strength material in the joint. Figure 1 verifies the assumption that the load bearing capacity of a heterogeneous configuration is supposed to equal the minimum strength of both homogenous assemblies. Material used in this study was a 1 mm low carbon equivalent Dual Phase 980 (DP980 LCE) steel.
Figure 1: Example of dissimilar configuration with CTS matching the “minimum rule”.B-6
The materials chosen for this study are in Table 1.
Table 1: Steel sheet samples.B-6
The material thickness combinations for all of the two sheet joints are shown in Table 2.
Table 2: Welded 2-sheet configurations.B-6
The three-sheet stackups all were made using the 1 mm DP980 LCE. These configurations were designed to understand what happens in such cases, knowing that three-sheet welding is very common in car body manufacturing. The three-sheet stackup configurations are shown in Figure 2 and are as were follows:
- a square DP980 coupon (patch) is inserted between the two classical cross-tension coupons for welding (1+patch+1 mm);
- two coupons oriented the same way welded with one coupon oriented in the transverse direction to form a cross-tension sample (1+[1+1] mm);
- same configuration as a) but the external coupon is removed by manual torsion before cross-tension testing (1+1+0 mm);
- same configuration as a), but the two coupons oriented the same way are first spot welded together strongly (with several spots) in the extremities, before the actual 3-sheet spot weld is done ([1++++1]+1 mm).
Figure 2: Three-sheet configurations based on 1mm DP980 LCE sample.B-6
The welding parameters for each configuration are listed in Table 3.
Table 3 : Welding parameters.B-6
CTS is strongly dependent on weld diameter (Figure 3).
Figure 3: Cross-tension Strength for TRIP800 configurations.B-6
CTS for the main DP980 configurations are shown as a function of weld diameter in Figures 4 and 5.
Figure 4: Cross-tension Strength for DP980 1+1, 1+2 and 2+2 configurations.B-6
Figure 5: Cross-tension Strength for DP980 1.25+1.25, 1.25+2 and 2+2 configurations.B-6
The three-sheet configurations based on 1mm DP980 LC results are shown in Figures 6 and 7. These results again verify that dissimilar configuration performances appear above the “minimum rule” assumption described in Figure 1.
Figure 6: Cross-tension Strength for DP980 1+1, 1+1+0 and 1+patch+1 configurations.B-6
![Figure 7: Cross-tension Strength for DP980 1+1, 1+2, 1+[1+1]and [1+++1]+1 configurations.](https://ahssinsights.org/wp-content/uploads/2020/07/3114c_Fig7.jpg)
Figure 7: Cross-tension Strength for DP980 1+1, 1+2, 1+[1+1]and [1+++1]+1 configurations.B-6
The observation that CTS is greater than predicted by the “minimum rule” has been called a “positive deviation” from the expected strengths.
This work concluded that while material qualification tests are frequently based on similar welding configurations, real car body applications are quite systematically dissimilar configurations. For spot welds failing in plug mode, the strength of the assembly only depends on the weakest material strength. In case of AHSS+AHSS welded combinations, however, things turn out to be different. Similar grade but dissimilar thickness High-Strength Steel configurations have been spot welded and tested in Cross-Tension. The following main conclusions can be highlighted:
- For dissimilar thickness configurations, the cross-tensile strength is above the standard “minimum rule” assumptions, this phenomenon being called a “positive deviation”;
- Limited thermal and notch location effects can explain part of this positive deviation, but the main reason is mechanical;
- As evidenced through several analytic and numerical studies, this mechanical effect is due to the less severe local stresses at the notch in case of uneven thickness, and improves the positive deviation when the thickness ratio increases. Although widely used for material qualification and scientific purposes, similar configurations appear as the worst case in terms of cross-tension performance for high strength steels. Actual vehicle design should consider positive deviation in dissimilar configurations to maximize the potential strength of spot welds in High-Strength steels.
RSW of Dissimilar Steel
This article is the summary of a paper entitled, “HAZ Softening of RSW of 3T Dissimilar Steel Stack-up”, Y. Lu., et al.L-15
Electromechanical Model
The study discusses the development of a 3D fully coupled thermo-electromechanical model for RSW of a three sheet (3T) stack-up of dissimilar steels. Figure 1 schematically shows the stack-up used in the study. The stack-up chosen is representative of the complex stack-ups used in BIW. Table 1 summarizes the nominal compositions of the three steels labeled in Figure 1.
Figure 1: Schematics of the 3T stack-up of 0.75-mm-thick JAC 270/1.4-mm-thick JSC 980/1.4-mm-thick JSC 590 steels.L-15
Table 1: Nominal Composition of Steels.L-15
JAC270 is a cold rolled Mild steel with a galvanneal coating having a minimum tensile strength of 270 MPa. JSC590 and JSC980 are bare cold rolled Dual Phase steels with a minimum tensile strength of 590 MPa and 980 MPa, respectively.
The electrodes used were CuZr dome-radius electrodes with a surface diameter of 6 mm. The welding parameters are listed in Table 2.
Table 2: Welding Parameters for Resistance Spot Welding of 3T Stack-Up of Steel Sheets.L-15
Figure 2 shows consistent nugget dimensions between simulation and experiment, supporting the validity of the RSW process model for 3T stack-up. The effect of welding current on nugget penetration into the thin sheet is similar to that on the nugget size. It increases rapidly at low welding current and saturates to 32% when the welding current is higher than 9 kA, as shown in Figure 2C.
Figure 2: Comparison between experimental and simulated results: A) Nugget geometry at 8 kA; B) nugget diameters; C) nugget penetration into the thin sheet as a function of welding current. In Figure 2A, the simulated nugget geometry is represented by the distribution of peak temperature (in Celsius). The two horizontal lines in Figure 2B represent the minimal nugget diameter at Interfaces A and B calculated, according to AWS D8.1M: 2007, Specification for Automotive Weld Quality Resistance Spot Welding of Steel. Due to limited number of samples available for testing, the variability in nugget dimensions at each welding current was not measuredL-15.
The results for nugget formation during RSW of the 3T stack-up are show in Figures 3-5. Figure 2 shows that, at the start of welding, the contact pressure at interface A (thin/thick) has a higher peak and drops more quickly along the radial direction than that at interface B (thick/thick). Due to the more localized contact area (Figure 3), a high current density can be observed at interface A, as shown in Figure 4A. Additionally, due to the high current density at interface A, localized heating is generated at this interface, as shown in Figure 5A.
Figure 3: Calculated contact pressure distribution at interfaces A (thin/thick) and B (thick/thick) at a welding time of 5 ms, current of 8 kA, and electrode force of 3.4 L-15
Figure 4: Calculated current density distribution at interfaces A (thin/thick) and B (thick/thick) at welding time of A — 5 ms; B — 200 and 300 ms.L-15
Figure 5: Temperature distribution during resistance spot welding of 3T stack-up at welding times of A) 5 ms; B) 102 ms; C) 300 ms. Welding current is 8 kA and electrode force is 3.4 kN. Calculated temperature is given in Celsius.L-15
As welding time increases, the contact area is expanded, resulting in a decrease of current density. The heat generation rate is shifted from interfaces to the bulk and the peak temperature occurs near the geometrical center of the stack-up.
Figure 6 illustrates that the predicted value corresponds well with the experimental data indicating a sound fitting to the isothermal tempering experimental data.
Figure 6: Comparison of the measured hardness with JMAK calculation showing the goodness of fit of the JSC 980 tempering kinetics parameters.L-15
Figure 7 shows the predicted hardness map of RSW 3T stack-up as well as the predicted and measured hardness profiles for JSC 980.
Figure 7: A) Predicted hardness map of resistance spot welded 3T stack-up; B) predicted and measured hardness profiles along the line marked in (A) for JSC 980.L-15
Automotive Welding Process Comparison
Introduction
A solution to improve the spot weld strength is to add a HS adhesive to the weld. Figure 1 illustrates the strength improvement obtained in static conditions when crash adhesive (example: Betamate 1496 Dow Automotive) is added. The trials are performed with 45-mm-wide and 16-mm adhesive bead samples.
Figure 1: TSS and CTS on DP 600.A-16
Another approach to improve the strength of welds is done by using laser welding instead of spot welding. The technologies based on remote welding optics have been introduced and a high productivity can be obtained. The effective welding time is maximized and a wide variety of weld geometries becomes feasible. Compared to spot welding, the main advantage of laser welding, regarding mechanical properties of the joint, is the possibility to adjust the weld dimension to the requirement. One may assume that, in tensile shear conditions, the weld strength depends linearly of the weld length (Figure 2).
Figure 2: Tensile-shear strength on laser weld stitches of different length.A-16
Comparing spot weld strength with laser weld strength cannot be restricted to the basic tensile shear test. Tests were performed to evaluate the weld strength in both quasi-static and dynamic conditions under different solicitations, on various UHSS combinations. The trials were performed on a high-speed testing machine, at 5 mm/min for the quasi-static tests and 0.5 m/s for the dynamic tests (pure shear, pure tear or mixed solicitation) (Figure 3). The strength at failure and the energy absorbed during the trial have been measured. It should be noticed that the energy absorbed depends also on the deformation of the sample. However, as all the trials were made according to the same sample geometry, the comparison of the results is relevant. Laser stitches were done with a 27-mm length. C- and S-shape welds were performed with the same overall weld length. This lead s to various apparent length and width of the welds. A shape factor, expressed as the ratio width/length of the weld, can be defined according to Table 1.
Figure 3: Sample geometry for quasi-static and dynamic tests. A-16
Table 1: Shape factor definition.A-16
The weld strength at failure can be easily described with an elliptic representation, with major axes representing pure shear and normal solicitation (Figure 4). For a reference spot weld corresponding to the upper limit of the weldability range, globally similar weld properties can be obtained with 27-mm laser welds. The spot weld equivalent length of 25-30 mm has been confirmed on other test cases on UHSS in the 1.5- to 2-mm range thickness. It has also been noticed that the spot weld equivalent length is shorter on thin mild steel (approximately 15-20 mm). This must be considered in case of shifting from spot to laser welding on a given structure. There is no major strain rate influence on the weld strength; the same order of magnitude is obtained in quasi-static and dynamic conditions.
Figure 4: Quasi-static and dynamic strength of welds, DP 600 2 mm+1.5 mm.A-16
The results in terms of energy absorbed by the sample are seen in Figure 5. In tearing conditions, both the strength at fracture and energy are lower for the spot weld than for the various laser welding procedures. In shear conditions, the strength at fracture is equivalent for all the welding processes. However, the energy absorption is more favorable to spot welds. This is due to the different fracture modes of the welds. IF fracture is observed on the laser welds under shearing solicitation (Figure 6). Even if the strength at failure is as high as for the spot weds, this brutal failure mode leads to lower total energy absorption.
Figure 5: Strength at fracture and energy absorption of HF1500P 1.8-mm + DP 600 1.5-mm samples for various welding conditions. A-16
Figure 6: IF fracture mode (left), “plug-out” fracture mode (right).A-16
Figure 7 represents the energy absorbed by omega-shaped structures and the corresponding number of welds that fail during the frontal crash test (here on TRIP 800 grade). It appears clearly that laser stitches have the highest rate of fracture during the crash test (33%). In standard spot welding, some weld fractures also occur. It is known that UHSS are more prone to partial IF fracture on coupons, and some welds fail as well during the crash test. By using either Weld-Bonding or adapted laser welding shape, there is no more weld fractures during the test, even if the parts are severely crashed and deformed. As a consequence, higher energy absorption is also observed.
Figure 7: Welding process and weld shape influence on the energy absorption and weld integrity on frontal crash tests. A-16
Regarding stiffness, up to 20% improvement can be obtained. The best results are obtained with continuous joints, and particularly using adhesives. Adhesive bonding and weld- bonding lead to the same results of the stiffness improvement only being due to the adhesive, not to the additional welds.
Figure 8 shows the evolution of the torsional stiffness with the joining process.
Figure 8: Evolution of the torsional stiffness with the joining process.A-16
Optimized laser joining design leads to same performances as a weld bonded sample regarding fracture modes seen in Figure 9.
Figure 9: Validation test case 1.2-mmTRIP 800/1.2-mm hat-shaped TRIP 800.
Top-hat crash boxes were tested across a range of AHSS materials including DP 1000. The spot weld’s energy absorbed increased linearly with increasing material strength. The adhesives were not suitable for crash applications as the adhesive peels open along the entire length of the joint. The welded bond samples perform much better than conventional spot welds. Across the entire range of materials there was a 20-30% increase in mean force when WB was used. The implications of such a large increase in crash performance are very significant. The results show that when a 600 MPa steel is weld bonded it can achieve the same crash performance as a 1000 MPa steel in spot-welded condition. It is also possible that some down gauging of materials could be achieved, but as the strength of the crash structure is highly dependent upon sheet thickness only small gage reductions would be possible.
Figure 10 shows the crash results for spot-welded and weld-bonded AHSS.
Figure 10: Crash results for spot-welded and weld-bonded AHSS.
There are numerous welding processes available for the welding of AHSS in automotive applications. Each of these processes has advantages and disadvantages that make them more or less applicable for particular applications. These qualities include joint efficiency, joint fit-up and design, joint strength, and stiffness, fracture mode, and cost effectiveness (equipment cost, production rates, etc.). The following data can allow for comparisons to be made for automotive application welding and joining processes, as well as possible repair substitutions.
Many tests were performed using lap and coach joints, reduced specimen overlap distance, and adjusted weld sizes to more closely represent typical joints consistent with automotive industry acceptance criteria. The tests were aimed at providing a baseline reference for a wide variety of welding and joining processes and material combinations. In general, there was no correlation between joint efficiency, normalized energy, and normalized stiffness. Joint efficiency was calculated by dividing the peak load of the joint by the peak load of the parent metal. Some processes, joint configurations and material combinations have high joint efficiency and energy, while others result in high joint efficiency but low energy. Few processes showed high values for all metrics across all materials and joint configurations (Figure 11). It was observed that peak loads tended to increase, on average, as material strength increased for lap joints (Figure 12). However, joint efficiency generally decreased as material strength increased. Therefore, joint strength did not increase in proportion to parent material strength increase for most of the processes and materials studied. Coach joints generally showed lower joint efficiency and stiffness than lap joints (Figure 13). Process and material combinations should be selected based on the required performance, joint design, and cost.A-12
Figure 11: Average peak loads (all processes combined).A-12
Figure 12: Lap shear average joint efficiency, normalized energy and stiffness (all processes combined).A-12
Figure 13: Coach peel average joint efficiency, normalized energy and stiffness (all processes combined).A-12
All Processes General Comparison
Numerous tests were performed using the most popular automotive joining processes including RSW, GMAW/brazing, laser welding/brazing, mechanical fasteners, and adhesive bonding. Joint efficiency and normalized energy of all the processes were compared for HSLA steels, DP 600 samples, DP 780 samples, and M190 samples. Joint efficiency was calculated as the peak load of the joint divided by the peak load of the parent metal. Energy was calculated as the area under the load/displacement curve up to peak load.
The materials used consisted of 1.2-mm EG HSLA, 1.2-mm galvanized DP 600, 1.0-mm GA DP 780, and 1.0-mm EG M190. The testing configuration matrix (Table 2) lists the materials and process combinations studied. The tolerance of weld lengths is ±10%. Lap-shear joints were centered in the overlap for all processes except lap fillet welds and brazes. Coach-peel joints were centered in the overlap for all processes (Figure 14). A-12
Table 2: Lap-shear (left) and coach-peel (right) test configuration matrix.A-12
Figure 14: Lap-shear and coach-peel set-up.A-12
DP 600 samples, DP 780 samples, and M190 samples. Joint efficiency was calculated as the peak load of the joint divided by the peak load of the parent metal. Energy was calculated as the area under the load/displacement curve up to peak load.
Self-piercing riveting with adhesive gave the greatest overall joint efficiency for the HSLA lap shear tests, while laser obtained the largest normalized energy (Figure 15a and 15b).
Figure 15a: Joint efficiency of HSLA lap-shear tests for all processes.A-12
Figure 15b: Normalized energy of HSLA lap-shear tests for all processes.A-12
Self-penetrating riveting with adhesive gave the greatest overall values for both joint efficiency and normalized energy for lap shear testing of DP 600 samples (Figure 16a and 16b). However, coach peel testing of DP 600 obtained the best results with laser welding (Figure 17).
Figure 17a: Joint efficiency of DP 600 lap shear for all processes.A-12
Figure 16b: Normalized energy of DP 600 lap shear for all processes.A-12
Figure 17: Joint efficiency and normalized energy of DP 600 coach peel for all processes.A-12
For the DP 780 lap-shear test, the best results out of all the tested processes were from laser/MIG welding, leading in both joint efficiency and normalized energy (Figure 18). Full laser welding produced the best results for coach-peel tests of the DP 780 samples (Figure 19).
Figure 18: Joint efficiency and normalized energy of DP 780 lap shear for all processes.A-12
Figure 19: Joint efficiency and normalized energy of DP 780 coach peel for all processes.A-12
The M190 lap-shear samples had the best joint efficiency using RSW with adhesive, but full laser welding gave better normalized energy (Figure 20). The coach peel tests also had the best normalized energy with full laser welding. The best joint efficiency of the coach peel tests was produced from laser welding with staples (Figure 21).
Figure 20: Joint efficiency and normalized energy of M190 lap shear for all processes.A-12
Figure 21: Joint efficiency and normalized energy of M190 coach peel for all processes.A-12
Cost Effectiveness Comparison: Spot Welding to Spot/Laser Welding Mixture
When automotive manufacturers are weighing the advantages and disadvantages of RSW to those of a spot/laser welding mixture process, cost effectiveness is a major concern. Spot/laser mixture welding has 38% lower operation cost compared to full spot welding because the laser installation performs more welds than a spot welding robot. Also, there are fewer robots to maintain and less consumables. The global cost is similar, but the spot/laser solution is about 4% less expensive overall. Figure 22 shows a cost comparison of spot welding and spot/laser welding.
Figure 22: Cost comparison of spot welding and spot/laser welding.A-16
GMAW Compared to Laser Welding
When comparing the advantages and disadvantages of GMAW to those of laser welding in automotive applications, joint efficiency is a key subject. Numerous welds were made using both processes on 15- and 25-mm-thick pieces of materials varying in strength. All results showed that laser welding continuously had greater joint efficiency than GMAW (Figure 23).
Figure 23: Joint efficiency of GMAW and laser welding for various steel strengths.
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![Figure 1. BS-SEM and EPMA Results of the Press-Hardened Blanks at 860°C [(a) 4 min, (b) 7 min, (c) 10 min, and (d) 900°C for 10 min (DL)].](http://ahssinsights.org/wp-content/uploads/2020/07/3112_Fig1.jpg)
![Figure 7: Cross-tension Strength for DP980 1+1, 1+2, 1+[1+1]and [1+++1]+1 configurations.](http://ahssinsights.org/wp-content/uploads/2020/07/3114c_Fig7.jpg)
