metallurgical bonding [4]. Joining by casting brings the advantages of hidden joint areas and the possibility for more complex geometries. In hybrid casting, molten metal is cast around an insert of metal in solid state [3]. A metallurgical bond with a shear strength of up to 7.7 MPa was achieved for a combination of aluminum and coated steel [5]. Joining by casting with an interlocking connection has been investigated for steel with aluminum [6] but is very common for polymers (joining by molding) as well [7]. In both cases, the process usually contains a surface structuring and an assembly step. For the surface structuring, processes like micromachining [8], micro-stamping [9], electron beam sculpturing [10], or additive processes with ball-head pins [6] or nanospikes [11] are possible.

To develop an efficient process chain for interlocking on large areas, this research focuses on a cold rolling process to form channel structures with undercuts onto the surface of sheet metal (Fig. 1 a) [12]. In an initial pass, a structured roller imprints a rectangular channel and rib structure in a steel sheet (Fig. 1 b). The undercuts are formed in a subsequent pass by flat rolling (Fig. 1 c). This is achieved as the tips of the ribs spread in lateral direction resulting in a dovetail geometry. For larger flattening, the channel side wall can fold over the channel bottom and form an inner notch [13]. This sheet material was inserted into a high-pressure die casting process at the Foundry Institute of RWTH Aachen University. The liquid aluminum flows into the undercut channels and interlocks as the melt solidifies [14]. The compound strength is achieved by a clamping between aluminum and steel and is therefore decisively influenced by the undercut width.

<figure><img src="image_1615.png" /><figcaption>Fig. 1: a) surface structuring by cold rolling, b) structural rolling and c) flattening</figcaption></figure>

In previous studies, hybrid casting tests with flat joining areas were performed, which achieved compound shear strengths up to 45 MPa [15]. As automotive components require an increased complexity of the parts, a section of an automotive roof cross beam was selected as a demonstrator part [16] (Fig. 2 c). Influencing factors of the die casting process have been analyzed and tensile strengths of up to 7.4 MPa were reached [17]. The complicated demonstrator geometry requires a bending of the structured sheet before casting (Fig. 2 a, b). It was assumed that the bending operation would affect the structure geometry. A narrowing of the channel opening would impede the filling with the aluminum melt and a reduction of the undercut width would reduce the compound strength. A previous study has revealed, that in the most unfavorable conditions investigated, the channel opening would decrease by 73% and the undercut width by 75% [18]. This study investigates how bending parameters like bending force or springback may depend on the imprinted surface structure. Similar to the previous study, the orientation of the structure to the
bending axis might also have an influence. This knowledge is of interest as precise contour geometries are needed, when inserting the bent sheet into a die casting mold. Therefore, this study aims to analyze how the surface structure affects the bending operations.

<figure><img src="image_1615.png" /><figcaption>Fig. 2: Process for hybrid components, a) and b) steel sheet insert with different surface structure orientations (0°, 90°), c) demonstrator part</figcaption></figure>

## Materials and Methods

Materials. The steel sheet inserts of the previous studies were produced of DC04, which is therefore used for this study as well. The material properties of DC04 are defined in Table 1 and Fig. 3 for the upcoming simulation. The flow curve was measured in stack layer compression tests up to a true strain of 0.3 at IBF and extrapolated with a modified Hollomon approach up to a true strain of 0.6 (Fig. 3). The extrapolation was based on the data points at higher strains to receive a good continuity between experimental and extrapolated data. As the bending force is considered in this study, the extrapolation might influence the accuracy of the results.

<table><caption>Table 1: Material properties of DC04 at room temperature</caption><thead><tr><td>Mass Density [t/mm³]</td><td>Young's Modulus [MPa]</td><td>Poisson's Ratio</td><td>σ₀ [MPa]</td></tr></thead><tbody><tr><td>7.85E-09</td><td>210,000</td><td>0.3</td><td>265.2</td></tr></tbody></table>

<figure><img src="image_1616.png" /><figcaption>Fig. 3: Flow curve of 2.0 mm DC04 at room temperature</figcaption></figure>

Structural rolling and flattening. Plates of DC04 were structured by cold rolling to create structured sheet material for the bending experiments. The plates were 2.0 mm thick, 200 mm wide and 245 mm long. The rolling was performed on a roll forming machine P3.160 by the company Dreistern GmbH & Co. KG with experimental parameters selected based on previous studies [13,15]. At the first rolling stand, a combination of an upper profiled roller (92.5 mm radius) and a lower flat roller (93.0 mm radius) were installed for the structuring. The profiled roller had seven
bending axis might also have an influence. This knowledge is of interest as precise contour geometries are needed, when inserting the bent sheet into a die casting mold. Therefore, this study aims to analyze how the surface structure affects the bending operations.

<figure><img src="image_1616.png" /><figcaption>Fig. 2: Process for hybrid components, a) and b) steel sheet insert with different surface structure orientations (0°, 90°), c) demonstrator part</figcaption></figure>

## Materials and Methods

Materials. The steel sheet inserts of the previous studies were produced of DC04, which is therefore used for this study as well. The material properties of DC04 are defined in Table 1 and Fig. 3 for the upcoming simulation. The flow curve was measured in stack layer compression tests up to a true strain of 0.3 at IBF and extrapolated with a modified Hollomon approach up to a true strain of 0.6 (Fig. 3). The extrapolation was based on the data points at higher strains to receive a good continuity between experimental and extrapolated data. As the bending force is considered in this study, the extrapolation might influence the accuracy of the results.

<table><caption>Table 1: Material properties of DC04 at room temperature</caption><thead><tr><td>Mass Density [t/mm³]</td><td>Young's Modulus [MPa]</td><td>Poisson's Ratio</td><td>σ₀ [MPa]</td></tr></thead><tbody><tr><td>7.85E-09</td><td>210,000</td><td>0.3</td><td>265.2</td></tr></tbody></table>

<figure><img src="image_1617.png" /><figcaption>Fig. 3: Flow curve of 2.0 mm DC04 at room temperature</figcaption></figure>

Structural rolling and flattening. Plates of DC04 were structured by cold rolling to create structured sheet material for the bending experiments. The plates were 2.0 mm thick, 200 mm wide and 245 mm long. The rolling was performed on a roll forming machine P3.160 by the company Dreistern GmbH & Co. KG with experimental parameters selected based on previous studies [13,15]. At the first rolling stand, a combination of an upper profiled roller (92.5 mm radius) and a lower flat roller (93.0 mm radius) were installed for the structuring. The profiled roller had seven
ribs and six channels, each 1.0 mm wide, resulting in 13 mm structured width. The milled channels were 0.5 mm deep and had a fillet radius of 0.05 mm at the rib edges. Two flat rollers (93.0 mm radius) were used on the second stand for the consecutive flattening pass. A reference sample was selected after structuring and flattening and the geometry was measured by cross-section preparation. A channel depth $\Delta h_s$ of approx. 465 µm resulted from the structuring. During flattening, the depth was reduced to $\Delta h_f = 343$ µm. For the simulation of the bending operation, an idealized structure geometry with the dimensions given in Table 2 was derived by averaging the measured values.

<table><caption>Table 2: Idealized reference structure geometry</caption><tr><td>channel opening width $w_{f-co}$</td><td>660 µm</td><td>rib width $w_{f-r}$</td><td>1346 µm</td></tr><tr><td>undercut width $w_{f-uc}$</td><td>57 µm</td><td>inner notch length $w_{f-no}$</td><td>93 µm</td></tr></table>

3-point bending. The bending test uses two support cylinders and a vertically moving punch with a load cell. The cylinders with a 5 mm radius are horizontally placed in a distance of 48 mm (center to center) and can freely rotate (Fig. 4). The structured sheet material from the previous cold rolling step was cut into specimens of 25 mm x 100 mm and placed on the support cylinders. The punch with a 5 mm radius moves vertically in the middle between the support cylinders to bend the sample. Depending on the displacement of the punch, a bending angle of up to 120° was reached.

<figure><img src="image_1617.png" /><figcaption>Fig. 4: Bending simulation with different surface structure orientation, a) 0°, in, b) 0°, out, c) 90°, in, d) 90°, out</figcaption></figure>

Simulation of the bending process. A 3D FE-model of the 3-point bending was set up in the software Abaqus (Fig. 4). The 3D approach was selected to model the bending behavior of the channels along (0°) and across (90°) the punch edge. Together with the channel structure on the inside (in) or outside (out) of the bend, four different configurations resulted (Fig. 4 a-d). As the alignment of the punch over a channel or a rib might affect the bending result, especially at small radii, the punch center was always positioned over the middle channel during the simulations. Two symmetry planes were used along the centerlines of the width and length of the specimen to reduce the computing effort. The sheet was given a fine mesh (approx. 0.05 mm) in areas of higher deformation like the bending apex to accurately display the formation of the structure. In areas of lower deformation such as the bending legs, a coarser mesh (approx. 1.0 mm) was selected to save computing effort. Furthermore, for the simulation of the channels across the bending edge (Fig. 4 c, d) the channels were truly modeled only around the bending apex. Thus, a section with the length of the punch diameter was given a surface structure. The bending legs, displayed by rectangular parts without surface structure, were tied to the structured section by a tie constraint. Hexagonal
ribs and six channels, each 1.0 mm wide, resulting in 13 mm structured width. The milled channels were 0.5 mm deep and had a fillet radius of 0.05 mm at the rib edges. Two flat rollers (93.0 mm radius) were used on the second stand for the consecutive flattening pass. A reference sample was selected after structuring and flattening and the geometry was measured by cross-section preparation. A channel depth $\Delta h_s$ of approx. 465 µm resulted from the structuring. During flattening, the depth was reduced to $\Delta h_f = 343$ µm. For the simulation of the bending operation, an idealized structure geometry with the dimensions given in Table 2 was derived by averaging the measured values.

<table><caption>Table 2: Idealized reference structure geometry</caption><tr><td>channel opening width $w_{f-co}$</td><td>660 µm</td><td>rib width $w_{f-r}$</td><td>1346 µm</td></tr><tr><td>undercut width $w_{f-uc}$</td><td>57 µm</td><td>inner notch length $w_{f-no}$</td><td>93 µm</td></tr></table>

3-point bending. The bending test uses two support cylinders and a vertically moving punch with a load cell. The cylinders with a 5 mm radius are horizontally placed in a distance of 48 mm (center to center) and can freely rotate (Fig. 4). The structured sheet material from the previous cold rolling step was cut into specimens of 25 mm x 100 mm and placed on the support cylinders. The punch with a 5 mm radius moves vertically in the middle between the support cylinders to bend the sample. Depending on the displacement of the punch, a bending angle of up to 120° was reached.

<figure><img src="image_1618.png" /><figcaption>Fig. 4: Bending simulation with different surface structure orientation, a) 0°, in, b) 0°, out, c) 90°, in, d) 90°, out</figcaption></figure>

Simulation of the bending process. A 3D FE-model of the 3-point bending was set up in the software Abaqus (Fig. 4). The 3D approach was selected to model the bending behavior of the channels along (0°) and across (90°) the punch edge. Together with the channel structure on the inside (in) or outside (out) of the bend, four different configurations resulted (Fig. 4 a-d). As the alignment of the punch over a channel or a rib might affect the bending result, especially at small radii, the punch center was always positioned over the middle channel during the simulations. Two symmetry planes were used along the centerlines of the width and length of the specimen to reduce the computing effort. The sheet was given a fine mesh (approx. 0.05 mm) in areas of higher deformation like the bending apex to accurately display the formation of the structure. In areas of lower deformation such as the bending legs, a coarser mesh (approx. 1.0 mm) was selected to save computing effort. Furthermore, for the simulation of the channels across the bending edge (Fig. 4 c, d) the channels were truly modeled only around the bending apex. Thus, a section with the length of the punch diameter was given a surface structure. The bending legs, displayed by rectangular parts without surface structure, were tied to the structured section by a tie constraint. Hexagonal
elements with reduced integration (C3D8R) were used for the mesh and a convergence study was performed regarding the fine-meshed areas. The punch and the support cylinders were assumed to be rigid surfaces. A reference point in the middle of the punch was used to extract the bending force. The contact between the sheet and the punch as well as the support cylinders was defined as *surface-to-surface* contact with a friction coefficient of 0.1 [19]. Higher friction coefficients of up to 0.3 were tested but this did not lead to any significant change in the bending force or springback. The process consists of a downward movement of the punch for the bending and an upward movement to relieve elastic deformation. In the simulation study, the punch displacement, the punch radii and the four orientations were varied. The punch displacement between 5 mm to 30 mm resulted in different bending angles. The simulations were validated with a punch radius of 5 mm. Further radii of 3 mm and 10 mm were simulated only.

## Results and discussion

The resulting bending angle as a function of the punch displacement for the bending radius R = 5 mm is presented in Fig. 5 a). For all four orientations, a very similar behavior is observed. The error between simulated and measured values is less than 4%. Only for the “0°, in” orientation, the simulation overestimates the experimental values by approx. 6%. However, it can be stated, that for all orientations the springback behavior is very similar, which is evidence for a good reproduction of the behavior.

<figure><img src="image_1618.png" /><figcaption>Fig. 5: a) resulting bending angle and b) maximum bending force for simulations and experiments with different surface structure orientations and a punch radius of 5 mm</figcaption></figure>

The maximum force needed for the bending operation is shown in Fig. 5 b). The force is initially increasing with increasing bending angle up to approx. 50° and is stagnating afterwards. The orientations with the channels along the bending edge (0°) show a good fit between simulations and experiments at approx. 0.78 kN in the steady area. For the 90° orientations, the simulations overestimate the experimental values with 0.98 kN by approx. 5%. A comparison of the 0° and 90° orientations shows an approx. 27% larger maximum bending force for channels across the bending edge. An orientation of the structure on the bending inside or outside radius showed no significant effect regarding the maximum bending force. Therefore, the analysis of the effect of the punch radius is exemplary presented for the structure on the inner radius.

Fig. 6 shows the influence of the punch radius on the maximum bending force. For the “0°, in” orientation, the bending force for radii 3 mm and 5 mm is very similar. However, for the 10 mm radius the force increases above 45° bending up to 0.86 kN, which is an additional 11%. A contrary effect is found for the “90°, in” orientation as the maximum bending force slightly decreases with larger radius.
elements with reduced integration (C3D8R) were used for the mesh and a convergence study was performed regarding the fine-meshed areas. The punch and the support cylinders were assumed to be rigid surfaces. A reference point in the middle of the punch was used to extract the bending force. The contact between the sheet and the punch as well as the support cylinders was defined as *surface-to-surface* contact with a friction coefficient of 0.1 [19]. Higher friction coefficients of up to 0.3 were tested but this did not lead to any significant change in the bending force or springback. The process consists of a downward movement of the punch for the bending and an upward movement to relieve elastic deformation. In the simulation study, the punch displacement, the punch radii and the four orientations were varied. The punch displacement between 5 mm to 30 mm resulted in different bending angles. The simulations were validated with a punch radius of 5 mm. Further radii of 3 mm and 10 mm were simulated only.

## Results and discussion

The resulting bending angle as a function of the punch displacement for the bending radius R = 5 mm is presented in Fig. 5 a). For all four orientations, a very similar behavior is observed. The error between simulated and measured values is less than 4%. Only for the “0°, in” orientation, the simulation overestimates the experimental values by approx. 6%. However, it can be stated, that for all orientations the springback behavior is very similar, which is evidence for a good reproduction of the behavior.

<figure><img src="image_1619.png" /><figcaption>Fig. 5: a) resulting bending angle and b) maximum bending force for simulations and experiments with different surface structure orientations and a punch radius of 5 mm</figcaption></figure>

The maximum force needed for the bending operation is shown in Fig. 5 b). The force is initially increasing with increasing bending angle up to approx. 50° and is stagnating afterwards. The orientations with the channels along the bending edge (0°) show a good fit between simulations and experiments at approx. 0.78 kN in the steady area. For the 90° orientations, the simulations overestimate the experimental values with 0.98 kN by approx. 5%. A comparison of the 0° and 90° orientations shows an approx. 27% larger maximum bending force for channels across the bending edge. An orientation of the structure on the bending inside or outside radius showed no significant effect regarding the maximum bending force. Therefore, the analysis of the effect of the punch radius is exemplary presented for the structure on the inner radius.

Fig. 6 shows the influence of the punch radius on the maximum bending force. For the “0°, in” orientation, the bending force for radii 3 mm and 5 mm is very similar. However, for the 10 mm radius the force increases above 45° bending up to 0.86 kN, which is an additional 11%. A contrary effect is found for the “90°, in” orientation as the maximum bending force slightly decreases with larger radius.
<figure><img src="image_1619.png" /><figcaption>Fig. 6: Maximum bending force for “0° and 90°, in” orientation at bending radii of 3 mm, 5 mm and 10 mm</figcaption></figure>

The effect of the punch radius on the springback is for the “0°, in” orientation (Fig. 7 a) minor. However, for the “90°, in” orientation, the 10 mm radius results in approx. 11% increased springback compared to 3 mm. This is a relevant difference to take into account for the process design.

<figure><img src="image_1620.png" /><figcaption>Fig. 7: Springback for 0° and 90°, in orientation at punch radii of 3 mm, 5 mm and 10 mm</figcaption></figure>

## Conclusion

The simulation and experimental values regarding the bending angle and maximum bending force for a 5 mm punch radius fit well. The springback seems to be influenced by the structure orientations only slightly. However, an increased springback behavior is found at larger bending radii. This is particular evident for the orientation across the bending edge on the inner radius (“90°, in” orientation) with an 11% increase.

The maximum bending force is influenced by the orientation of the channels across or along the bending edge of the punch. Channels across the bending edge results in approx. 27% larger force and the force is reduced by increasing bending radius. Contrary to that, the force increases when using larger punch radii for orientations with channels along the punch edge. An orientation of the structure on the bending inside or outside radius showed no significant effect. However, this knowledge will help with the design of structured and bent parts. For further study, it might be of
<figure><img src="image_1620.png" /><figcaption>Fig. 6: Maximum bending force for “0° and 90°, in” orientation at bending radii of 3 mm, 5 mm and 10 mm</figcaption></figure>

The effect of the punch radius on the springback is for the “0°, in” orientation (Fig. 7 a) minor. However, for the “90°, in” orientation, the 10 mm radius results in approx. 11% increased springback compared to 3 mm. This is a relevant difference to take into account for the process design.

<figure><img src="image_1621.png" /><figcaption>Fig. 7: Springback for 0° and 90°, in orientation at punch radii of 3 mm, 5 mm and 10 mm</figcaption></figure>

## Conclusion

The simulation and experimental values regarding the bending angle and maximum bending force for a 5 mm punch radius fit well. The springback seems to be influenced by the structure orientations only slightly. However, an increased springback behavior is found at larger bending radii. This is particular evident for the orientation across the bending edge on the inner radius (“90°, in” orientation) with an 11% increase.

The maximum bending force is influenced by the orientation of the channels across or along the bending edge of the punch. Channels across the bending edge results in approx. 27% larger force and the force is reduced by increasing bending radius. Contrary to that, the force increases when using larger punch radii for orientations with channels along the punch edge. An orientation of the structure on the bending inside or outside radius showed no significant effect. However, this knowledge will help with the design of structured and bent parts. For further study, it might be of
interest to investigate how the alignment of the punch above a rib or channel influences the bending behavior.

**Conflict of Interest**

The authors declare that they have no conflict of interest.

**Acknowledgement**

This study was funded by the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation) - Project-ID 245566034.

**References**

[1] K. Mori, Y. Abe, A review on mechanical joining of aluminium and high strength steel sheets by plastic deformation. International Journal of Lightweight Materials and Manufacture 1 (1) (2018) 1-11. https://doi.org/10.1016/j.ijlmm.2018.02.002

[2] K. Mori, N. Bay, L. Fratini, F. Micari, A.E. Tekkaya, Joining by plastic deformation. CIRP Annals - Manufacturing Technology 62 (2) (2013) 673-694. https://doi.org/10.1016/j.cirp.2013.05.004

[3] T. Noguchi, K. Asano, S. Hiratsuka, H. Miyahara, Trends of composite casting technology and joining technology for castings in Japan. International Journal of Cast Metals Research 21 (1-4) (2008) 219-225. https://doi.org/10.1179/136404608X361981

[4] M. Salamati, M. Soltanpour, A. Fazli, A. Zajkani, Processing and tooling considerations in joining by forming technologies; part A—mechanical joining. Int J Adv Manuf Technol 101 (1-4) (2019) 261-315. https://doi.org/10.1007/s00170-018-2823-y

[5] X. Fang, J. Gundlach, J.-J. Schipperges, X. Jiang, On the Steel-Aluminum Hybrid Casting by Sand Casting. J. Mater. Eng. Perform. 27 (12) (2018) 6415-6425. https://doi.org/10.1007/s11665-018-3717-8

[6] S. Ucsnik, R. Gradinger, A. Becirovic, A. Waldhör, Enhanced Performance of Steel-Aluminium Cast Nodes through Cold Metal Transfer. Mater. Sci. Forum 765 (2013) 736-740. https://doi.org/10.4028/www.scientific.net/MSF.765.736

[7] E.E. Feistauer, J.F. dos Santos, S.T. Amancio-Filho, A review on direct assembly of through-the-thickness reinforced metal-polymer composite hybrid structures. Polymer Engineering & Sci 59 (4) (2019) 661-674. https://doi.org/10.1002/pen.25022

[8] Y. Ma, G. Zhang, S. Cao, Z. Huo, J. Han, S. Ma, Z. Huang, A Review of Advances in Fabrication Methods and Assistive Technologies of Micro-Structured Surfaces. Processes 11 (5) (2023) 1337. https://doi.org/10.3390/pr11051337

[9] S. Müller, M. Brand, K. Dröder, D. Meiners, Increasing the Structural Integrity of Hybrid Plastics-Metal Parts by an Innovative Mechanical Interlocking Effect. Materials Science Forum 825-826 (2015) 417-424. https://doi.org/10.4028/www.scientific.net/MSF.825-826.417

[10] A.L. Buxton, A. Ferhati, R.J. Glen, B.G.I. Dance, D. Mullen, T. Karayiannis, EB Surface Engineering for High Performance Heat Exchangers, First International Electron Beam Welding Conference (2009).

[11] H. Abe, J.C. Chung, T. Mori, A. Hosoi, K.M. Jespersen, H. Kawada, The effect of nanospike structures on direct bonding strength properties between aluminum and carbon fiber reinforced thermoplastics. Composites Part B: Engineering 172 (2019) 26-32. https://doi.org/10.1016/j.compositesb.2019.05.025
interest to investigate how the alignment of the punch above a rib or channel influences the bending behavior.

**Conflict of Interest**

The authors declare that they have no conflict of interest.

**Acknowledgement**

This study was funded by the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation) - Project-ID 245566034.

**References**

[1] K. Mori, Y. Abe, A review on mechanical joining of aluminium and high strength steel sheets by plastic deformation. International Journal of Lightweight Materials and Manufacture 1 (1) (2018) 1-11. https://doi.org/10.1016/j.ijlmm.2018.02.002

[2] K. Mori, N. Bay, L. Fratini, F. Micari, A.E. Tekkaya, Joining by plastic deformation. CIRP Annals - Manufacturing Technology 62 (2) (2013) 673-694. https://doi.org/10.1016/j.cirp.2013.05.004

[3] T. Noguchi, K. Asano, S. Hiratsuka, H. Miyahara, Trends of composite casting technology and joining technology for castings in Japan. International Journal of Cast Metals Research 21 (1-4) (2008) 219-225. https://doi.org/10.1179/136404608X361981

[4] M. Salamati, M. Soltanpour, A. Fazli, A. Zajkani, Processing and tooling considerations in joining by forming technologies; part A—mechanical joining. Int J Adv Manuf Technol 101 (1-4) (2019) 261-315. https://doi.org/10.1007/s00170-018-2823-y

[5] X. Fang, J. Gundlach, J.-J. Schipperges, X. Jiang, On the Steel-Aluminum Hybrid Casting by Sand Casting. J. Mater. Eng. Perform. 27 (12) (2018) 6415-6425. https://doi.org/10.1007/s11665-018-3717-8

[6] S. Ucsnik, R. Gradinger, A. Becirovic, A. Waldhör, Enhanced Performance of Steel-Aluminium Cast Nodes through Cold Metal Transfer. Mater. Sci. Forum 765 (2013) 736-740. https://doi.org/10.4028/www.scientific.net/MSF.765.736

[7] E.E. Feistauer, J.F. dos Santos, S.T. Amancio-Filho, A review on direct assembly of through-the-thickness reinforced metal-polymer composite hybrid structures. Polymer Engineering & Sci 59 (4) (2019) 661-674. https://doi.org/10.1002/pen.25022

[8] Y. Ma, G. Zhang, S. Cao, Z. Huo, J. Han, S. Ma, Z. Huang, A Review of Advances in Fabrication Methods and Assistive Technologies of Micro-Structured Surfaces. Processes 11 (5) (2023) 1337. https://doi.org/10.3390/pr11051337

[9] S. Müller, M. Brand, K. Dröder, D. Meiners, Increasing the Structural Integrity of Hybrid Plastics-Metal Parts by an Innovative Mechanical Interlocking Effect. Materials Science Forum 825-826 (2015) 417-424. https://doi.org/10.4028/www.scientific.net/MSF.825-826.417

[10] A.L. Buxton, A. Ferhati, R.J. Glen, B.G.I. Dance, D. Mullen, T. Karayiannis, EB Surface Engineering for High Performance Heat Exchangers, First International Electron Beam Welding Conference (2009).

[11] H. Abe, J.C. Chung, T. Mori, A. Hosoi, K.M. Jespersen, H. Kawada, The effect of nanospike structures on direct bonding strength properties between aluminum and carbon fiber reinforced thermoplastics. Composites Part B: Engineering 172 (2019) 26-32. https://doi.org/10.1016/j.compositesb.2019.05.025
[12] S. Senge, J. Brachmann, G. Hirt, A. Bührig-Polaczek, Interlocking Multi-Material Components made of Structured Steel Sheets and High-Pressure Die Cast Aluminium, ESAFORM (2017), 190007. https://doi.org/10.1063/1.5008220

[13] A. Ringel, J. Lohmar, Optimization of the Surface Geometry in Structured Cold Rolling for Interlocking of Formed and Die-Cast Metal Components. DDF 414 (2022) 89-94. https://doi.org/10.4028/p-z54p05

[14] J. Jakumeit, H. Behnken, R. Laqua, J. Eiken, J. Brachmann, Multi-scale simulation of hybrid light metal structures produced by high pressure die casting. IOP Conf. Ser.: Mater. Sci. Eng. 861 (2020) 12035. https://doi.org/10.1088/1757-899X/861/1/012035

[15] S. Senge, J. Brachmann, G. Hirt, A. Bührig-Polaczek, Evaluation of interlocking bond strength between structured 1.0338 steel sheets and high pressure die cast AlMg5Si2, in: Livan Fratini, Rosa Di Lorenzo, Gianluca Buffa, Guiseppe Ingaro, (Ed.), ESAFORM (2018), 40019. https://doi.org/10.1063/1.5034873

[16] D. Joop, Präzisionsbestimmende Faktoren bei der Herstellung blechverstärkter Hybridstrukturen im Druckguss: Grenzflächenanalyse und Verzugsoptimierung, Gießerei-Institut, RWTH Aachen University (2018). https://doi.org/10.18154/RWTH-2019-00954

[17] L. Bruckmeier, A. Ringel, U. Vroomen, D. Bailly, A. Bührig-Polaczek, Influence of High-Pressure Die Casting Process Parameters on the Compound Strength of Hybrid Components with Undercut Sheet Metal. Metals 13 (10) (2023) 1717. https://doi.org/10.3390/met13101717

[18] A. Ringel, G. Hirt, Bending behavior of structured steel sheets with undercuts for interlocking with Al die-cast metal. Int J Mater Form 17 (1) (2024). https://doi.org/10.1007/s12289-023-01797-6

[19] A. Mousavi, K.S. Ridolfi, A. Brosius, Influence of alternating bending on springback behavior of parts in deep drawing process with macro-structured tools. MATEC Web Conf. 190 (2018) 14005. https://doi.org/10.1051/matecconf/201819014005
[12] S. Senge, J. Brachmann, G. Hirt, A. Bührig-Polaczek, Interlocking Multi-Material Components made of Structured Steel Sheets and High-Pressure Die Cast Aluminium, ESAFORM (2017), 190007. https://doi.org/10.1063/1.5008220

[13] A. Ringel, J. Lohmar, Optimization of the Surface Geometry in Structured Cold Rolling for Interlocking of Formed and Die-Cast Metal Components. DDF 414 (2022) 89-94. https://doi.org/10.4028/p-z54p05

[14] J. Jakumeit, H. Behnken, R. Laqua, J. Eiken, J. Brachmann, Multi-scale simulation of hybrid light metal structures produced by high pressure die casting. IOP Conf. Ser.: Mater. Sci. Eng. 861 (2020) 12035. https://doi.org/10.1088/1757-899X/861/1/012035

[15] S. Senge, J. Brachmann, G. Hirt, A. Bührig-Polaczek, Evaluation of interlocking bond strength between structured 1.0338 steel sheets and high pressure die cast AlMg5Si2, in: Livan Fratini, Rosa Di Lorenzo, Gianluca Buffa, Guiseppe Ingaro, (Ed.), ESAFORM (2018), 40019. https://doi.org/10.1063/1.5034873

[16] D. Joop, Präzisionsbestimmende Faktoren bei der Herstellung blechverstärkter Hybridstrukturen im Druckguss: Grenzflächenanalyse und Verzugsoptimierung, Gießerei-Institut, RWTH Aachen University (2018). https://doi.org/10.18154/RWTH-2019-00954

[17] L. Bruckmeier, A. Ringel, U. Vroomen, D. Bailly, A. Bührig-Polaczek, Influence of High-Pressure Die Casting Process Parameters on the Compound Strength of Hybrid Components with Undercut Sheet Metal. Metals 13 (10) (2023) 1717. https://doi.org/10.3390/met13101717

[18] A. Ringel, G. Hirt, Bending behavior of structured steel sheets with undercuts for interlocking with Al die-cast metal. Int J Mater Form 17 (1) (2024). https://doi.org/10.1007/s12289-023-01797-6

[19] A. Mousavi, K.S. Ridolfi, A. Brosius, Influence of alternating bending on springback behavior of parts in deep drawing process with macro-structured tools. MATEC Web Conf. 190 (2018) 14005. https://doi.org/10.1051/matecconf/201819014005
# In-situ CT of the clinching process – Influence of settling effects due to process interruptions

KÖHLER Daniel<sup>a*</sup>, KUPFER Robert<sup>b</sup>, TROSCHITZ Juliane<sup>c</sup> and GUDE Maik<sup>d</sup>

Institute of Lightweight Engineering and Polymer Technology, TUD Dresden University of Technology, Germany

<sup>a</sup>daniel.koehler3@tu-dresden.de, <sup>b</sup>robert.kupfer@tu-dresden.de, <sup>c</sup>juliane.troschitz@tu-dresden.de, <sup>d</sup>maik.gude@tu-dresden.de

**Keywords:** Clinching, Testing, Computed Tomography, In-Situ CT

**Abstract.** In lightweight constructions, clinching represents a cost-effective solution, in which joints are produced by local cold forming of the joining parts. Clinching phenomena are typically evaluated using destructive testing methods. While these methods influence the clinch point's state, in-situ computed tomography (in-situ CT) is able to explore the clinching process with a specimen under load. Here, the path-controlled clinching process is interrupted at certain displacement levels and the specimen is scanned by CT while remaining in a stationary state. These interruptions are always accompanied by settling effects reducing the reaction force. Therefore, in this work, the influence of these interruptions on the force-displacement behavior during clinching and on the final clinch point's geometric properties is investigated.

## Introduction

Clinching technology. The increased interest in lightweight components for car body constructions requires joining methods, which are able to join different materials. Likewise, in light of ever-increasing legal requirements on material recycling [1], joining solutions that facilitate material separation with low contamination [2] are gaining increasing interest. In terms of joining lightweight materials such as aluminum, clinching poses a good alternative to bolting, riveting or bonding with high strength adhesives as it allows a good recyclability. This is achieved because a force and form fit joint is created by locally deforming two join partners without requiring any auxiliary joining elements [3]. Moreover, different materials can be joined [4], and the costs can be reduced to 30 - 60 % compared to resistance spot welding [5].

Designing clinch points. The design of clinch points is described in DVS/EFB 3420 [6], providing reference values for quasi-static shear strength based on individual material-thickness combinations. However, in practice, joint design often relies on experience, iterative testing, and tool revisions [7]. This results in a tool configuration which is specific to a single material-thickness combination often leading to an unnecessary material utilization. To enhance the design and adapt clinch tools more efficiently to varying conditions, a better understanding of occurring phenomena in the clinch point is necessary. Besides that, a more precise investigation of the phenomena can also facilitate an accurate validation of numerical models which are also required for the design process.

Classical characterization methods. Intuitively, phenomena occurring during the clinching process are visually or optically inspected after manufacturing [8]. This way, asymmetries and peripheral cracks can be found, however phenomena in depth remain unknown. Using destructive metallographic examination according to [8], deformation phenomena within the clinch point can be visualized. For this, the specimen is cut close to the cross-section of interest, then it is embedded in resin, ground, and optionally polished. A macroscopic analysis of the cross section enables the determination of geometric parameters such as undercut and neck thickness, as well as asymmetries and cracks within the cutting plane. Additionally, this method reveals the interface

<figure><img src="image_1621.png" /></figure>

Content from this work may be used under the terms of the Creative Commons Attribution 3.0 license. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under license by Materials Research Forum LLC.
# In-situ CT of the clinching process – Influence of settling effects due to process interruptions

KÖHLER Daniel<sup>a*</sup>, KUPFER Robert<sup>b</sup>, TROSCHITZ Juliane<sup>c</sup> and GUDE Maik<sup>d</sup>

Institute of Lightweight Engineering and Polymer Technology, TUD Dresden University of Technology, Germany

<sup>a</sup>daniel.koehler3@tu-dresden.de, <sup>b</sup>robert.kupfer@tu-dresden.de, <sup>c</sup>juliane.troschitz@tu-dresden.de, <sup>d</sup>maik.gude@tu-dresden.de

**Keywords:** Clinching, Testing, Computed Tomography, In-Situ CT

**Abstract.** In lightweight constructions, clinching represents a cost-effective solution, in which joints are produced by local cold forming of the joining parts. Clinching phenomena are typically evaluated using destructive testing methods. While these methods influence the clinch point's state, in-situ computed tomography (in-situ CT) is able to explore the clinching process with a specimen under load. Here, the path-controlled clinching process is interrupted at certain displacement levels and the specimen is scanned by CT while remaining in a stationary state. These interruptions are always accompanied by settling effects reducing the reaction force. Therefore, in this work, the influence of these interruptions on the force-displacement behavior during clinching and on the final clinch point's geometric properties is investigated.

**Introduction**

Clinching technology. The increased interest in lightweight components for car body constructions requires joining methods, which are able to join different materials. Likewise, in light of ever-increasing legal requirements on material recycling [1], joining solutions that facilitate material separation with low contamination [2] are gaining increasing interest. In terms of joining lightweight materials such as aluminum, clinching poses a good alternative to bolting, riveting or bonding with high strength adhesives as it allows a good recyclability. This is achieved because a force and form fit joint is created by locally deforming two join partners without requiring any auxiliary joining elements [3]. Moreover, different materials can be joined [4], and the costs can be reduced to 30 - 60 % compared to resistance spot welding [5].

Designing clinch points. The design of clinch points is described in DVS/EFB 3420 [6], providing reference values for quasi-static shear strength based on individual material-thickness combinations. However, in practice, joint design often relies on experience, iterative testing, and tool revisions [7]. This results in a tool configuration which is specific to a single material-thickness combination often leading to an unnecessary material utilization. To enhance the design and adapt clinch tools more efficiently to varying conditions, a better understanding of occurring phenomena in the clinch point is necessary. Besides that, a more precise investigation of the phenomena can also facilitate an accurate validation of numerical models which are also required for the design process.

Classical characterization methods. Intuitively, phenomena occurring during the clinching process are visually or optically inspected after manufacturing [8]. This way, asymmetries and peripheral cracks can be found, however phenomena in depth remain unknown. Using destructive metallographic examination according to [8], deformation phenomena within the clinch point can be visualized. For this, the specimen is cut close to the cross-section of interest, then it is embedded in resin, ground, and optionally polished. A macroscopic analysis of the cross section enables the determination of geometric parameters such as undercut and neck thickness, as well as asymmetries and cracks within the cutting plane. Additionally, this method reveals the interface

<figure><img src="image_1622.png" /></figure>
and outer edges of the clinch point in the respective cross-section, which is often utilized for validating numerical models [9]. In order to understand the evolution of phenomena, clinch points can be tested and analyzed in step-setting tests [10]. Here, a series of metallographic examinations is conducted with clinch points which were manufactured at different stages of the clinching process. However, this way, phenomena such as closing cracks and springback effects influence the investigated displacement state. Furthermore, investigations at stages when no sufficient undercut is created are hardly possible. But most importantly, the investigation of the evolution of particular phenomena is restrained due to material and manufacturing variations.

In-situ CT. As an alternative, X-ray in-situ computed tomography (in-situ CT) allows the three-dimensional visualization of clinching phenomena under load. Hence, the evolution of individual phenomena can be investigated. For this, a CT system is combined with a testing machine mounted with CT suitable clinching tools. Applying this method using a lab-based CT system, the specimen is clinched stepwise in increasing displacement levels. Usually, at each displacement level, the testing process is interrupted, and the specimen is rested in order to allow a CT scan of the specimen's displacement state. These interruptions are accompanied by settling effects reducing the reaction force measured by the testing machine over the resting time. Optionally, X-ray in-situ CT with synchrotron X-ray source allows scanning time of the order of seconds or below [11]. This could enable CT investigations of continuous clinching processes, though at a significantly decreased manufacturing speed (common clinching speed: 2 mm/s). Because of the high scanning costs of synchrotron CT, lab-based X-ray in-situ CT remain a popular alternative and shall be investigated in this work.

The lab-based in-situ CT method has been gaining increased interest over the last decade [12]. It was already applied to investigating aluminum foams under tensile loading [13] or metal ceramic composites in compression tests [14]. In terms of joints, shear tests of adhesively bonded riveted lap joints [15] or the pull out behavior of inserts out of carbon fiber reinforced plastic [16] were already investigated by in-situ CT. Moreover, the in-situ CT of the manufacturing and testing process of clinch points was introduced in [17] and [18], respectively. However, to the author's knowledge, there are no published investigations on the influence of the interruptions, which are inherent for lab-based in-situ CT, on the clinching process or comparable high deformation processes. So far, it is usually assumed that the settling effects have no significant impact on the process of interest. In order to clarify on this potential impact a more detailed investigation is required.

Hence, in this work, two set of specimens are investigated. First, the specimens are clinched continuously (Conti), second the specimens are clinched with longer interruptions (StopGo) at several displacement levels. The force-displacement behavior is analyzed and the geometric properties of the final clinch point is compared statistically.

## Materials and Methods

Experimental setup. For the experiments, clinch points made of two 2 mm thick aluminum round sheets (ø 40 mm) are produced in two different manufacturing procedures in the same in-situ CT clinching setup. For reference, continuous clinching is conducted for ten specimens with a cross beam displacement rate of 2 mm/s (Conti specimen set). Here the punch is moved to a total displacement. After that the punch is instantly withdrawn resulting in a bottom thickness of 0.70±0.01. Another ten specimens (StopGo specimen set) are clinched using the same total punch displacement. To represent the in-situ CT process, the punch movement is interrupted and held in position for 15 min at distinctive displacement levels in different phases of the clinching process. It was interrupted at 25 % (offsetting phase), 50 % (upsetting phase), 75 % (upsetting phase) and 100 % (flow pressing phase) of the total punch displacement (stop points). A preload of 1.7 kN is chosen to align the experiments at a force occurring shortly after the punch touched the sheets.
and outer edges of the clinch point in the respective cross-section, which is often utilized for validating numerical models [9]. In order to understand the evolution of phenomena, clinch points can be tested and analyzed in step-setting tests [10]. Here, a series of metallographic examinations is conducted with clinch points which were manufactured at different stages of the clinching process. However, this way, phenomena such as closing cracks and springback effects influence the investigated displacement state. Furthermore, investigations at stages when no sufficient undercut is created are hardly possible. But most importantly, the investigation of the evolution of particular phenomena is restrained due to material and manufacturing variations.

In-situ CT. As an alternative, X-ray in-situ computed tomography (in-situ CT) allows the three-dimensional visualization of clinching phenomena under load. Hence, the evolution of individual phenomena can be investigated. For this, a CT system is combined with a testing machine mounted with CT suitable clinching tools. Applying this method using a lab-based CT system, the specimen is clinched stepwise in increasing displacement levels. Usually, at each displacement level, the testing process is interrupted, and the specimen is rested in order to allow a CT scan of the specimen's displacement state. These interruptions are accompanied by settling effects reducing the reaction force measured by the testing machine over the resting time. Optionally, X-ray in-situ CT with synchrotron X-ray source allows scanning time of the order of seconds or below [11]. This could enable CT investigations of continuous clinching processes, though at a significantly decreased manufacturing speed (common clinching speed: 2 mm/s). Because of the high scanning costs of synchrotron CT, lab-based X-ray in-situ CT remain a popular alternative and shall be investigated in this work.

The lab-based in-situ CT method has been gaining increased interest over the last decade [12]. It was already applied to investigating aluminum foams under tensile loading [13] or metal ceramic composites in compression tests [14]. In terms of joints, shear tests of adhesively bonded riveted lap joints [15] or the pull out behavior of inserts out of carbon fiber reinforced plastic [16] were already investigated by in-situ CT. Moreover, the in-situ CT of the manufacturing and testing process of clinch points was introduced in [17] and [18], respectively. However, to the author's knowledge, there are no published investigations on the influence of the interruptions, which are inherent for lab-based in-situ CT, on the clinching process or comparable high deformation processes. So far, it is usually assumed that the settling effects have no significant impact on the process of interest. In order to clarify on this potential impact a more detailed investigation is required.

Hence, in this work, two set of specimens are investigated. First, the specimens are clinched continuously (Conti), second the specimens are clinched with longer interruptions (StopGo) at several displacement levels. The force-displacement behavior is analyzed and the geometric properties of the final clinch point is compared statistically.

## Materials and Methods

Experimental setup. For the experiments, clinch points made of two 2 mm thick aluminum round sheets (ø 40 mm) are produced in two different manufacturing procedures in the same in-situ CT clinching setup. For reference, continuous clinching is conducted for ten specimens with a cross beam displacement rate of 2 mm/s (Conti specimen set). Here the punch is moved to a total displacement. After that the punch is instantly withdrawn resulting in a bottom thickness of 0.70±0.01. Another ten specimens (StopGo specimen set) are clinched using the same total punch displacement. To represent the in-situ CT process, the punch movement is interrupted and held in position for 15 min at distinctive displacement levels in different phases of the clinching process. It was interrupted at 25 % (offsetting phase), 50 % (upsetting phase), 75 % (upsetting phase) and 100 % (flow pressing phase) of the total punch displacement (stop points). A preload of 1.7 kN is chosen to align the experiments at a force occurring shortly after the punch touched the sheets.
Evaluation of the force displacement behavior. First, it is tested how well the force-displacement curves of both manufacturing procedures align. This is done through statistically testing for a significant deviation. For this, the force, the normalized standard deviation of the forces, and the normalized differences between the means of the forces of both procedures are investigated over the displacement. Comparing the forces, clear deviations at the stop points are expected due to settling effects in the StopGo tests. Therefore, it is investigated how well the forces of both procedures align shortly after the stop point. This is done by testing the deviation between both procedures statistically for significance at a displacement of 0.2 mm after each stop point (evaluation point). Additionally, the maximum forces are evaluated statistically. For comparing both manufacturing procedures, both mean values (t-test) and variances (F-test) of the investigated forces at the evaluation points and the maximum forces are tested for homogeneity. Thus, it shall be verified that the force displacement curves of both procedures do not deviate from each other significantly.

Evaluation of the final geometry. Furthermore, it is verified whether the final geometries of the respective specimen sets align. This is done by statistically testing whether the specimen sets' geometry deviate from each other significantly. For this purpose, the clinch points are analyzed using (ex-situ) CT (cf. scanning parameters in Table 1). Here, a stack of six specimens mounted in a specimen holder is scanned using the CT system V|TOME|X L450¹ <fnref n="2" /> with a helix scanning trajectory. After the scans are reconstructed using the Feldkamp algorithm in phoenix datos|x 2 Reconstruction<fnref n="3" />, the volumes are edited in VG Studio Max 3.4⁴. For the evaluation, the specimens' surfaces are to be compared. The surface of each specimen is determined by first creating a contour based on the iso 50 value and excluding wrongly detected voids and noise particles. Then, this contour is used as a starting contour to determine the maximum grey value gradient within an area of 10 µm, which is used as the surface to be analyzed.

For the evaluation, the specimens' surfaces are to be compared. The surface of each specimen is determined by first creating a contour based on the iso 50 value and excluding wrongly detected voids and noise particles. Then, this contour is used as a starting contour to determine the maximum grey value gradient within an area of 10 µm, which is used as the surface to be analyzed. For comparing the contours, the bottom thickness is measured along the rotation axis of a fitted cone (cf. Figure 1). The neck thicknesses are measured horizontally (cf. translated surface in Figure 1) in at a constant distance of 2.45 mm to a fitted surface on the die faced sheet surface (cf. fitted surface in Figure 1)) in the cross section planes 0° and 90° relative to the material texture of both sheets<fnref n="5" />. The neck thicknesses are then evaluated for the 0° and 90° cross section planes respectively.

The neck thicknesses are then evaluated for the 0° and 90° cross section planes respectively. For verifying that the final geometries do not deviate significantly, the mean values (t-test) and variances (F-test) of the neck and bottom thicknesses for both specimen sets are evaluated.

<footnote n="1">¹ GE Sensing & Inspection Technologies GmbH</footnote>

<footnote n="2">² 300 kV microfocus, flat detector with pixel size 200 x 200 µm, 2016 x 2016 pixels, 16 bit</footnote>

<footnote n="3">³ GE Sensing & Inspection Technologies GmbH</footnote>

<footnote n="4">⁴ Volume Graphics GmbH, Heidelberg, Deutschland</footnote>

<footnote n="5">⁵ Prior clinching the sheets are positioned towards each other with aligned texture orientation</footnote>
Evaluation of the force displacement behavior. First, it is tested how well the force-displacement curves of both manufacturing procedures align. This is done through statistically testing for a significant deviation. For this, the force, the normalized standard deviation of the forces, and the normalized differences between the means of the forces of both procedures are investigated over the displacement. Comparing the forces, clear deviations at the stop points are expected due to settling effects in the StopGo tests. Therefore, it is investigated how well the forces of both procedures align shortly after the stop point. This is done by testing the deviation between both procedures statistically for significance at a displacement of 0.2 mm after each stop point (evaluation point). Additionally, the maximum forces are evaluated statistically. For comparing both manufacturing procedures, both mean values (t-test) and variances (F-test) of the investigated forces at the evaluation points and the maximum forces are tested for homogeneity. Thus, it shall be verified that the force displacement curves of both procedures do not deviate from each other significantly.

Evaluation of the final geometry. Furthermore, it is verified whether the final geometries of the respective specimen sets align. This is done by statistically testing whether the specimen sets' geometry deviate from each other significantly. For this purpose, the clinch points are analyzed using (ex-situ) CT (cf. scanning parameters in Table 1). Here, a stack of six specimens mounted in a specimen holder is scanned using the CT system V|TOME|X L450¹ <fnref n="2" /> with a helix scanning trajectory. After the scans are reconstructed using the Feldkamp algorithm in phoenix datos|x 2 Reconstruction<fnref n="3" />, the volumes are edited in VG Studio Max 3.4⁴. For the evaluation, the specimens' surfaces are to be compared. The surface of each specimen is determined by first creating a contour based on the iso 50 value and excluding wrongly detected voids and noise particles. Then, this contour is used as a starting contour to determine the maximum grey value gradient within an area of 10 µm, which is used as the surface to be analyzed.

For the evaluation, the specimens' surfaces are to be compared. The surface of each specimen is determined by first creating a contour based on the iso 50 value and excluding wrongly detected voids and noise particles. Then, this contour is used as a starting contour to determine the maximum grey value gradient within an area of 10 µm, which is used as the surface to be analyzed. For comparing the contours, the bottom thickness is measured along the rotation axis of a fitted cone (cf. Figure 1). The neck thicknesses are measured horizontally (cf. translated surface in Figure 1) in at a constant distance of 2.45 mm to a fitted surface on the die faced sheet surface (cf. fitted surface in Figure 1)) in the cross section planes 0° and 90° relative to the material texture of both sheets<fnref n="5" />. The neck thicknesses are then evaluated for the 0° and 90° cross section planes respectively.

The neck thicknesses are then evaluated for the 0° and 90° cross section planes respectively. For verifying that the final geometries do not deviate significantly, the mean values (t-test) and variances (F-test) of the neck and bottom thicknesses for both specimen sets are evaluated.

<footnote n="1">¹ GE Sensing & Inspection Technologies GmbH</footnote>

<footnote n="2">² 300 kV microfocus, flat detector with pixel size 200 x 200 µm, 2016 x 2016 pixels, 16 bit</footnote>

<footnote n="3">³ GE Sensing & Inspection Technologies GmbH</footnote>

<footnote n="4">⁴ Volume Graphics GmbH, Heidelberg, Deutschland</footnote>

<footnote n="5">⁵ Prior clinching the sheets are positioned towards each other with aligned texture orientation</footnote>
<table><caption>Table 1: Applied parameters for the CT measurement and image analysis</caption><thead><tr><td>Parameter</td><td>Unit</td><td>Value</td></tr></thead><tbody><tr><td>Acceleration voltage</td><td>[kV]</td><td>150</td></tr><tr><td>Tube current</td><td>[µA]</td><td>80</td></tr><tr><td>X-ray projections</td><td></td><td>2016</td></tr><tr><td>Exposure time</td><td>[ms]</td><td>500</td></tr><tr><td>Voxel size</td><td>[µm]</td><td>10</td></tr><tr><td>Skip</td><td></td><td>1</td></tr><tr><td>Averaging</td><td></td><td>5</td></tr><tr><td>Physical filter</td><td>[mm]</td><td>1 (Cu)</td></tr><tr><td>Focal spot size</td><td>[µm]</td><td>12</td></tr></tbody></table>

<figure><img src="image_1622.png" /><figcaption>Figure 1: 0° cross section of the clinch point CT volume in 2D (a) and 3D (b) with the fitted geometries supporting the measurement of geometric properties in the final clinch point.</figcaption></figure>

## Results and Discussion

Force-displacement evaluation. The force-displacement diagrams for both sample variations are shown in Figure 1. The stop points, evaluation points, and close-up views of the graphs at these points can be seen. A notable observation is the significant deviation of one curve, C04_A_CV_440⁶, from the StopGo tests. Aside from that, after each stop point the forces of both procedures align well, though the StopGo graphs show a decreasing alignment to the Conti graphs with increasing displacement level. Evaluating Figure 2 supports the identification of specimen C04_A_CV_440 as an outlier (boxplot method). Therefore, this specimen is excluded from the statistical analysis. In the boxplots, it can be seen that the variances and the mean of both specimen sets seem to be similar at all evaluation points.

<footnote n="6">⁶ Unified specimen notation in Collaborative Research Center TRR285</footnote>
<table><caption>Table 1: Applied parameters for the CT measurement and image analysis</caption><thead><tr><td>Parameter</td><td>Unit</td><td>Value</td></tr></thead><tbody><tr><td>Acceleration voltage</td><td>[kV]</td><td>150</td></tr><tr><td>Tube current</td><td>[µA]</td><td>80</td></tr><tr><td>X-ray projections</td><td></td><td>2016</td></tr><tr><td>Exposure time</td><td>[ms]</td><td>500</td></tr><tr><td>Voxel size</td><td>[µm]</td><td>10</td></tr><tr><td>Skip</td><td></td><td>1</td></tr><tr><td>Averaging</td><td></td><td>5</td></tr><tr><td>Physical filter</td><td>[mm]</td><td>1 (Cu)</td></tr><tr><td>Focal spot size</td><td>[µm]</td><td>12</td></tr></tbody></table>

<figure><img src="image_1623.png" /><figcaption>Figure 1: 0° cross section of the clinch point CT volume in 2D (a) and 3D (b) with the fitted geometries supporting the measurement of geometric properties in the final clinch point.</figcaption></figure>

## Results and Discussion

Force-displacement evaluation. The force-displacement diagrams for both sample variations are shown in Figure 1. The stop points, evaluation points, and close-up views of the graphs at these points can be seen. A notable observation is the significant deviation of one curve, C04_A_CV_440⁶, from the StopGo tests. Aside from that, after each stop point the forces of both procedures align well, though the StopGo graphs show a decreasing alignment to the Conti graphs with increasing displacement level. Evaluating Figure 2 supports the identification of specimen C04_A_CV_440 as an outlier (boxplot method). Therefore, this specimen is excluded from the statistical analysis. In the boxplots, it can be seen that the variances and the mean of both specimen sets seem to be similar at all evaluation points.

<footnote n="6">⁶ Unified specimen notation in Collaborative Research Center TRR285</footnote>
<figure><img src="image_1623.png" /><figcaption>Figure 2: Force-displacement diagram of Conti and StopGo tests with the stop points and the evaluated points</figcaption></figure>
<figure><img src="image_1624.png" /><figcaption>Figure 2: Force-displacement diagram of Conti and StopGo tests with the stop points and the evaluated points</figcaption></figure>
<figure><img src="image_1624.png" /><figcaption>Figure 3: Box plot diagrams at the evaluation points and $F_{max}$</figcaption></figure>

In Figure 4 a, the normalized standard deviations of both manufacturing procedures over the displacement are displayed. Before the first stop point, no difference can be seen. After the first stop point an increasing variance is notable for the StopGo procedure. However, near the maximum force both variances are similar again. In light of the maximum normalized standard deviation of 1 %, it seems that both procedures exhibit similarly low deviations over the whole process. The forces at the evaluation points were checked for normal distribution (Shapiro-Wilk and Anderson-Darling test) and the F-test was conducted (test for variance homogeneity). In Table 2 the results for each evaluation point are shown. In a one-tailed test, it is tested whether StopGo tests exhibit a significantly higher variance as Conti tests. It can be concluded that the variance is not significantly different (min p-value 0.2312 > 0.05). Consequently, no significant impact of the interruptions on the variance of the forces can be found.

To get an impression of the mean value homogeneity, the averaged force at each displacement is calculated resulting in a respresentive force-displacement plot for each specimen set. The normalized<fnref n="7" /> difference between the averaged StopGo and Conti graphs gives an impression of the mean value homogeneity (cf. Figure 3 b). It can be seen that the normalized mean force difference before the first stop point is below 0.7 %. The stop points cause a high difference in the mean value due to the drop in force caused by settling effects. After the stop point, the differences in the mean value decrease. However, the rate of decline is slowing down after each stop point. Shortly before the maximum force, the difference reduces to 0.5 %. Overall, the differences are at a relatively low level. Conducting the t-test (test for mean homogeneity) at the evaluation points reveals a

<footnote n="7">⁷ using the averaged Conti graph</footnote>
<figure><img src="image_1625.png" /><figcaption>Figure 3: Box plot diagrams at the evaluation points and $F_{max}$</figcaption></figure>

In Figure 4 a, the normalized standard deviations of both manufacturing procedures over the displacement are displayed. Before the first stop point, no difference can be seen. After the first stop point an increasing variance is notable for the StopGo procedure. However, near the maximum force both variances are similar again. In light of the maximum normalized standard deviation of 1 %, it seems that both procedures exhibit similarly low deviations over the whole process. The forces at the evaluation points were checked for normal distribution (Shapiro-Wilk and Anderson-Darling test) and the F-test was conducted (test for variance homogeneity). In Table 2 the results for each evaluation point are shown. In a one-tailed test, it is tested whether StopGo tests exhibit a significantly higher variance as Conti tests. It can be concluded that the variance is not significantly different (min p-value 0.2312 > 0.05). Consequently, no significant impact of the interruptions on the variance of the forces can be found.

To get an impression of the mean value homogeneity, the averaged force at each displacement is calculated resulting in a respresentive force-displacement plot for each specimen set. The normalized<fnref n="7" /> difference between the averaged StopGo and Conti graphs gives an impression of the mean value homogeneity (cf. Figure 3 b). It can be seen that the normalized mean force difference before the first stop point is below 0.7 %. The stop points cause a high difference in the mean value due to the drop in force caused by settling effects. After the stop point, the differences in the mean value decrease. However, the rate of decline is slowing down after each stop point. Shortly before the maximum force, the difference reduces to 0.5 %. Overall, the differences are at a relatively low level. Conducting the t-test (test for mean homogeneity) at the evaluation points reveals a

<footnote n="7">⁷ using the averaged Conti graph</footnote>
decreasing p-value with increasing displacement. A significant deviation (α = 0.05) is found for the evaluation point $s_{max_{75\%}} + 0.2$ mm and $F_{max}$.

<figure><img src="image_1625.png" /><figcaption>Figure 4: The normalized standard deviation of both procedures over the displacement (a). Normalized differences of the means of the forces of both procedures over the displacement (b). Outlier excluded.</figcaption></figure>

<table><caption>Table 2: F- and t-test results for the evaluation of the force-displacement behavior (outlier excluded)</caption><thead><tr><th>Test method</th><th>Evaluated point</th><th>Statistic</th><th>p-value</th><th>Significant deviation (p-values > 0.05)</th></tr></thead><tbody><tr><td rowspan="4">F-test</td><td>s<sub>max<sub>25%</sub></sub> + 0.2 mm</td><td>1.58</td><td>0.2649</td><td>Wrong</td></tr><tr><td>s<sub>max<sub>50%</sub></sub> + 0.2 mm</td><td>1.64</td><td>0.2476</td><td>Wrong</td></tr><tr><td>s<sub>max<sub>75%</sub></sub> + 0.2 mm</td><td>1.70</td><td>0.2312</td><td>Wrong</td></tr><tr><td>F<sub>max</sub></td><td>0.958</td><td>0.5291</td><td>Wrong</td></tr><tr><td rowspan="4">t-test</td><td>s<sub>max<sub>25%</sub></sub> + 0.2 mm</td><td>-0.59</td><td>0.5600</td><td>Wrong</td></tr><tr><td>s<sub>max<sub>50%</sub></sub> + 0.2 mm</td><td>-1.88</td><td>0.0773</td><td>Wrong</td></tr><tr><td>s<sub>max<sub>75%</sub></sub> + 0.2 mm</td><td>-2.71</td><td>0.0146</td><td>True</td></tr><tr><td>F<sub>max</sub></td><td>-2.86</td><td>0.0108</td><td>True</td></tr></tbody></table>

Geometry evaluation. The box plots for the investigated geometric properties are shown in Figure 4. The bottom thickness clearly deviates between both specimen sets. At the neck thickness, no obvious deviations can be seen in between the specimen sets. However, a clear difference can be seen when comparing the results from the 0° and 90° cross section plane for both specimen sets. Evaluating the F- and t-test, it can be seen that all investigated properties have no significant variance inhomogeneity. However, the mean values of the bottom thickness deviate significantly in between both specimen sets.
decreasing p-value with increasing displacement. A significant deviation (α = 0.05) is found for the evaluation point $s_{max_{75\%}} + 0.2 \text{ mm}$ and $F_{max}$.

<figure><img src="image_1626.png" /><figcaption>Figure 4: The normalized standard deviation of both procedures over the displacement (a). Normalized differences of the means of the forces of both procedures over the displacement (b). Outlier excluded.</figcaption></figure>

<table><caption>Table 2: F- and t-test results for the evaluation of the force-displacement behavior (outlier excluded)</caption><thead><tr><th>Test method</th><th>Evaluated point</th><th>Statistic</th><th>p-value</th><th>Significant deviation (p-values > 0.05)</th></tr></thead><tbody><tr><td rowspan="4">F-test</td><td>s<sub>max<sub>25\%</sub></sub> + 0.2 mm</td><td>1.58</td><td>0.2649</td><td>Wrong</td></tr><tr><td>s<sub>max<sub>50\%</sub></sub> + 0.2 mm</td><td>1.64</td><td>0.2476</td><td>Wrong</td></tr><tr><td>s<sub>max<sub>75\%</sub></sub> + 0.2 mm</td><td>1.70</td><td>0.2312</td><td>Wrong</td></tr><tr><td>F<sub>max</sub></td><td>0.958</td><td>0.5291</td><td>Wrong</td></tr><tr><td rowspan="4">t-test</td><td>s<sub>max<sub>25\%</sub></sub> + 0.2 mm</td><td>-0.59</td><td>0.5600</td><td>Wrong</td></tr><tr><td>s<sub>max<sub>50\%</sub></sub> + 0.2 mm</td><td>-1.88</td><td>0.0773</td><td>Wrong</td></tr><tr><td>s<sub>max<sub>75\%</sub></sub> + 0.2 mm</td><td>-2.71</td><td>0.0146</td><td>True</td></tr><tr><td>F<sub>max</sub></td><td>-2.86</td><td>0.0108</td><td>True</td></tr></tbody></table>

Geometry evaluation. The box plots for the investigated geometric properties are shown in Figure 4. The bottom thickness clearly deviates between both specimen sets. At the neck thickness, no obvious deviations can be seen in between the specimen sets. However, a clear difference can be seen when comparing the results from the 0° and 90° cross section plane for both specimen sets. Evaluating the F- and t-test, it can be seen that all investigated properties have no significant variance inhomogeneity. However, the mean values of the bottom thickness deviate significantly in between both specimen sets.
<figure><img src="image_1626.png" /><figcaption>Figure 5: Measured bottom thickness (a) and neck thicknesses in the respective cross section planes (b) for both specimen sets.</figcaption></figure>

<table><caption>Table 3: F- and t-test results for the evaluation of the geometry (outlier excluded)</caption><thead><tr><th>Test method</th><th>Evaluated point</th><th>Statistic</th><th>p-value</th><th>Significant deviation (p-values > 0.05)</th></tr></thead><tbody><tr><td rowspan="3">F-test</td><td>Bottom thickness x<sub>b</sub></td><td>0.43</td><td>0.8798</td><td>Wrong</td></tr><tr><td>Neck thickness x<sub>b</sub> 0°</td><td>0.77</td><td>0.7078</td><td>Wrong</td></tr><tr><td>Neck thickness x<sub>b</sub> 90°</td><td>0.47</td><td>0.9383</td><td>Wrong</td></tr><tr><td rowspan="3">t-test</td><td>Bottom thickness x<sub>b</sub></td><td>5.44</td><td>4.35E-05</td><td>True</td></tr><tr><td>Neck thickness x<sub>b</sub> 0°</td><td>0.51</td><td>0.6064</td><td>Wrong</td></tr><tr><td>Neck thickness x<sub>b</sub> 90°</td><td>0.80</td><td>0.4244</td><td>Wrong</td></tr></tbody></table>

## Summary and Conclusions

Investigating the deformation phenomena during clinching processes can preferably be done using in-situ CT as this method does not influence the process compared to destructive methods. When conducting a lab-based in-situ CT investigation, the clinch point is created stepwise. At displacement levels of interest, the process is interrupted and the specimen is scanned by CT while remaining in a stationary state. These interruptions are always accompanied by reductions of the reaction force caused by settling effects. While this investigation method is already applied to clinching and also to other processes, the effect of interrupting the process of interest is not investigated in detail yet. Therefore, in this work, the clinching forces of interrupted and continuous clinching processes as well as the resulting geometry of the clinch point are compared statistically. While the variances of these properties of both clinching procedures are well aligned, the mean values at higher clinching forces and of the bottom thicknesses deviated significantly. At the stepwise processes the mean clinching force is higher and the mean bottom thickness is smaller. However, the differences are relatively low. The means and the variances of the neck thicknesses of both specimen sets did not deviate significantly.
<figure><img src="image_1627.png" /><figcaption>Figure 5: Measured bottom thickness (a) and neck thicknesses in the respective cross section planes (b) for both specimen sets.</figcaption></figure>

<table><caption>Table 3: F- and t-test results for the evaluation of the geometry (outlier excluded)</caption><thead><tr><th>Test method</th><th>Evaluated point</th><th>Statistic</th><th>p-value</th><th>Significant deviation (p-values > 0.05)</th></tr></thead><tbody><tr><td rowspan="3">F-test</td><td>Bottom thickness x<sub>b</sub></td><td>0.43</td><td>0.8798</td><td>Wrong</td></tr><tr><td>Neck thickness x<sub>b</sub> 0°</td><td>0.77</td><td>0.7078</td><td>Wrong</td></tr><tr><td>Neck thickness x<sub>b</sub> 90°</td><td>0.47</td><td>0.9383</td><td>Wrong</td></tr><tr><td rowspan="3">t-test</td><td>Bottom thickness x<sub>b</sub></td><td>5.44</td><td>4.35E-05</td><td>True</td></tr><tr><td>Neck thickness x<sub>b</sub> 0°</td><td>0.51</td><td>0.6064</td><td>Wrong</td></tr><tr><td>Neck thickness x<sub>b</sub> 90°</td><td>0.80</td><td>0.4244</td><td>Wrong</td></tr></tbody></table>

## Summary and Conclusions

Investigating the deformation phenomena during clinching processes can preferably be done using in-situ CT as this method does not influence the process compared to destructive methods. When conducting a lab-based in-situ CT investigation, the clinch point is created stepwise. At displacement levels of interest, the process is interrupted and the specimen is scanned by CT while remaining in a stationary state. These interruptions are always accompanied by reductions of the reaction force caused by settling effects. While this investigation method is already applied to clinching and also to other processes, the effect of interrupting the process of interest is not investigated in detail yet. Therefore, in this work, the clinching forces of interrupted and continuous clinching processes as well as the resulting geometry of the clinch point are compared statistically. While the variances of these properties of both clinching procedures are well aligned, the mean values at higher clinching forces and of the bottom thicknesses deviated significantly. At the stepwise processes the mean clinching force is higher and the mean bottom thickness is smaller. However, the differences are relatively low. The means and the variances of the neck thicknesses of both specimen sets did not deviate significantly.
It can be concluded that performing a clinching process with interruptions as it is common in in-situ CT investigations has a relatively small but partly significant impact on the resulting clinching forces and geometric properties of the clinch point. At the continuous tests, the punch is directly withdrawn after reaching the final punch displacement in contrast to the stepwise test. Thus, the differences in the bottom thickness can be explained by settling effects during the interruption at the last stop point leaving a higher plastic deformation.

Considering the small differences in reaction forces and geometric properties it can be concluded that an in-situ CT investigation can sufficiently accurately represent the continuous clinching process. Thus, with this method, important phenomena and the evolvement of the final geometry can be analyzed. For a further alignment of the geometric properties, particularly the bottom thickness, the in-situ CT clinching process needs to be adapted, potentially by reducing the total punch displacement.

## References

[1] European Commission, "A European Strategy for Plastics in a Circular Economy: Communication from the Commission to the European Parliament, the Council, the European Economic and Social Committee and the Committee of the Regions," Brussels, Jan. 2018. Accessed: Mar. 20 2022. [Online]. Available: https://ec.europa.eu/environment/circular-economy/pdf/plastics-strategy.pdf

[2] K. Ragaert, L. Delva, and K. van Geem, "Mechanical and chemical recycling of solid plastic waste," *Waste management (New York, N.Y.)*, vol. 69, pp. 24–58, 2017. https://doi.org/10.1016/j.wasman.2017.07.044

[3] V. K. Soo, P. Compston, and M. Doolan, "The influence of joint technologies on ELV recyclability," *Waste management (New York, N.Y.)*, vol. 68, pp. 421–433, 2017. https://doi.org/10.1016/j.wasman.2017.07.020.

[4] X. He, "Clinching for sheet materials," *Science and technology of advanced materials*, vol. 18, no. 1, pp. 381–405, 2017. https://doi.org/10.1080/14686996.2017.1320930

[5] P. Kovács and M. Tisza, "Investigation of clinch joints made of similar and dissimilar materials," *IOP Conf. Ser.: Mater. Sci. Eng.*, vol. 426, p. 12028, 2018. https://doi.org/10.1088/1757-899X/426/1/012028

[6] *Technical Bulletin DVS 3420: Prüfung der Eigenschaften mechanisch und kombiniert mittels Kleben gefertigter Verbindungen*, 3420, Working Group V10 "Mechanical Joining" by the joint committee for DVS and EFB, Düsseldorf, Jun. 2021.

[7] G. Meschut and F. F. Menne, *Automatisierte Variantenreduzierung durch virtuelle Verbindungsauslegung beim Halbhohlstanznieten*, 1st ed. Aachen: Shaker Verlag, 2018.

[8] *Testing of properties of joints - Testing of properties of mechanical and hybrid (mechanical/bonded) joints: Prüfung der Eigenschaften mechanisch und kombiniert mittels Kleben gefertigter Verbindungen*, 3480-1, Working Group V10 "Mechanical Joining" by the joint committee for DVS and EFB, Düsseldorf, Dec. 2007.

[9] S. Coppieters, S. Cooreman, P. Lava, H. Sol, P. van Houtte, and D. Debruyne, "Reproducing the experimental pull-out and shear strength of clinched sheet metal connections using FEA," *Int J Mater Form*, vol. 4, no. 4, pp. 429–440, 2011. https://doi.org/10.1007/s12289-010-1023-6

[10] R. Kupfer et al., "Clinching of Aluminum Materials – Methods for the Continuous Characterization of Process, Microstructure and Properties," *Journal of Advanced Joining Processes*, vol. 5, p. 100108, 2022. https://doi.org/10.1016/j.jajp.2022.100108
It can be concluded that performing a clinching process with interruptions as it is common in in-situ CT investigations has a relatively small but partly significant impact on the resulting clinching forces and geometric properties of the clinch point. At the continuous tests, the punch is directly withdrawn after reaching the final punch displacement in contrast to the stepwise test. Thus, the differences in the bottom thickness can be explained by settling effects during the interruption at the last stop point leaving a higher plastic deformation.

Considering the small differences in reaction forces and geometric properties it can be concluded that an in-situ CT investigation can sufficiently accurately represent the continuous clinching process. Thus, with this method, important phenomena and the evolvement of the final geometry can be analyzed. For a further alignment of the geometric properties, particularly the bottom thickness, the in-situ CT clinching process needs to be adapted, potentially by reducing the total punch displacement.

## References

[1] European Commission, "A European Strategy for Plastics in a Circular Economy: Communication from the Commission to the European Parliament, the Council, the European Economic and Social Committee and the Committee of the Regions," Brussels, Jan. 2018. Accessed: Mar. 20 2022. [Online]. Available: https://ec.europa.eu/environment/circular-economy/pdf/plastics-strategy.pdf

[2] K. Ragaert, L. Delva, and K. van Geem, "Mechanical and chemical recycling of solid plastic waste," *Waste management (New York, N.Y.)*, vol. 69, pp. 24–58, 2017. https://doi.org/10.1016/j.wasman.2017.07.044

[3] V. K. Soo, P. Compston, and M. Doolan, "The influence of joint technologies on ELV recyclability," *Waste management (New York, N.Y.)*, vol. 68, pp. 421–433, 2017. https://doi.org/10.1016/j.wasman.2017.07.020.

[4] X. He, "Clinching for sheet materials," *Science and technology of advanced materials*, vol. 18, no. 1, pp. 381–405, 2017. https://doi.org/10.1080/14686996.2017.1320930

[5] P. Kovács and M. Tisza, "Investigation of clinch joints made of similar and dissimilar materials," *IOP Conf. Ser.: Mater. Sci. Eng.*, vol. 426, p. 12028, 2018. https://doi.org/10.1088/1757-899X/426/1/012028

[6] *Technical Bulletin DVS 3420: Prüfung der Eigenschaften mechanisch und kombiniert mittels Kleben gefertigter Verbindungen*, 3420, Working Group V10 "Mechanical Joining" by the joint committee for DVS and EFB, Düsseldorf, Jun. 2021.

[7] G. Meschut and F. F. Menne, *Automatisierte Variantenreduzierung durch virtuelle Verbindungsauslegung beim Halbhohlstanznieten*, 1st ed. Aachen: Shaker Verlag, 2018.

[8] *Testing of properties of joints - Testing of properties of mechanical and hybrid (mechanical/bonded) joints: Prüfung der Eigenschaften mechanisch und kombiniert mittels Kleben gefertigter Verbindungen*, 3480-1, Working Group V10 "Mechanical Joining" by the joint committee for DVS and EFB, Düsseldorf, Dec. 2007.

[9] S. Coppieters, S. Cooreman, P. Lava, H. Sol, P. van Houtte, and D. Debruyne, "Reproducing the experimental pull-out and shear strength of clinched sheet metal connections using FEA," *Int J Mater Form*, vol. 4, no. 4, pp. 429–440, 2011. https://doi.org/10.1007/s12289-010-1023-6

[10] R. Kupfer et al., "Clinching of Aluminum Materials – Methods for the Continuous Characterization of Process, Microstructure and Properties," *Journal of Advanced Joining Processes*, vol. 5, p. 100108, 2022. https://doi.org/10.1016/j.jajp.2022.100108
[11] J. Dewanckele, M. A. Boone, F. Coppens, D. van Loo, and A. P. Merkle, "Innovations in laboratory-based dynamic micro-CT to accelerate in situ research," *Journal of microscopy*, vol. 277, no. 3, pp. 197-209, 2020. https://doi.org/10.1111/jmi.12879

[12] E. A. Zwanenburg, M. A. Williams, and J. M. Warnett, "Review of high-speed imaging with lab-based x-ray computed tomography," *Meas. Sci. Technol.*, vol. 33, no. 1, p. 12003, 2022. https://doi.org/10.1088/1361-6501/ac354a

[13] Y. Amani, S. Dancette, E. Maire, J. Adrien, and J. Lachambre, "Two-Scale Tomography Based Finite Element Modeling of Plasticity and Damage in Aluminum Foams," *Materials (Basel, Switzerland)*, vol. 11, no. 10, 2018. https://doi.org/10.3390/ma11101984

[14] J. Schukraft, C. Lohr, and K. A. Weidenmann, "2D and 3D in-situ mechanical testing of an interpenetrating metal ceramic composite consisting of a slurry-based ceramic foam and AlSi10Mg," *Composite Structures*, vol. 263, p. 113742, 2021. https://doi.org/10.1016/j.compstruct.2021.113742

[15] R. Füßel, M. Gude, and A. Mertel, "In-situ X-ray computed tomography analysis of adhesively bonded riveted lap joints," in *17th European Conference on Composite Materials*, Munich, 2016.

[16] F. Pottmeyer, J. Bittner, P. Pinter, and K. A. Weidenmann, "In-Situ CT Damage Analysis of Metal Inserts Embedded in Carbon Fiber-Reinforced Plastics," *Exp Mech*, vol. 57, no. 9, pp. 1411-1422, 2017. https://doi.org/10.1007/s11340-017-0312-0

[17] D. Köhler, R. Kupfer, J. Troschitz, and M. Gude, "Clinching in In Situ CT—A Novel Validation Method for Mechanical Joining Processes," in *The Minerals, Metals & Materials Series, NUMISHEET 2022: Proceedings of the 12th international conference and workshop on*, K. Inal, J. Levesque, M. Worswick, and C. Butcher, Eds., [S.l.]: SPRINGER, 2022, pp. 833-840.

[18] D. Köhler, R. Kupfer, J. Troschitz, and M. Gude, "In Situ Computed Tomography-Analysis of a Single-Lap Shear Test with Clinch Points," *Materials (Basel, Switzerland)*, vol. 14, no. 8, 2021. https://doi.org/10.3390/ma14081859
[11] J. Dewanckele, M. A. Boone, F. Coppens, D. van Loo, and A. P. Merkle, "Innovations in laboratory-based dynamic micro-CT to accelerate in situ research," *Journal of microscopy*, vol. 277, no. 3, pp. 197-209, 2020. https://doi.org/10.1111/jmi.12879

[12] E. A. Zwanenburg, M. A. Williams, and J. M. Warnett, "Review of high-speed imaging with lab-based x-ray computed tomography," *Meas. Sci. Technol.*, vol. 33, no. 1, p. 12003, 2022. https://doi.org/10.1088/1361-6501/ac354a

[13] Y. Amani, S. Dancette, E. Maire, J. Adrien, and J. Lachambre, "Two-Scale Tomography Based Finite Element Modeling of Plasticity and Damage in Aluminum Foams," *Materials (Basel, Switzerland)*, vol. 11, no. 10, 2018. https://doi.org/10.3390/ma11101984

[14] J. Schukraft, C. Lohr, and K. A. Weidenmann, "2D and 3D in-situ mechanical testing of an interpenetrating metal ceramic composite consisting of a slurry-based ceramic foam and AlSi10Mg," *Composite Structures*, vol. 263, p. 113742, 2021. https://doi.org/10.1016/j.compstruct.2021.113742

[15] R. Füßel, M. Gude, and A. Mertel, "In-situ X-ray computed tomography analysis of adhesively bonded riveted lap joints," in *17th European Conference on Composite Materials*, Munich, 2016.

[16] F. Pottmeyer, J. Bittner, P. Pinter, and K. A. Weidenmann, "In-Situ CT Damage Analysis of Metal Inserts Embedded in Carbon Fiber-Reinforced Plastics," *Exp Mech*, vol. 57, no. 9, pp. 1411-1422, 2017. https://doi.org/10.1007/s11340-017-0312-0

[17] D. Köhler, R. Kupfer, J. Troschitz, and M. Gude, "Clinching in In Situ CT—A Novel Validation Method for Mechanical Joining Processes," in *The Minerals, Metals & Materials Series, NUMISHEET 2022: Proceedings of the 12th international conference and workshop on*, K. Inal, J. Levesque, M. Worswick, and C. Butcher, Eds., [S.l.]: SPRINGER, 2022, pp. 833-840.

[18] D. Köhler, R. Kupfer, J. Troschitz, and M. Gude, "In Situ Computed Tomography-Analysis of a Single-Lap Shear Test with Clinch Points," *Materials (Basel, Switzerland)*, vol. 14, no. 8, 2021. https://doi.org/10.3390/ma14081859
# Flexible self-pierce riveting and clinching with a single joining system using the same unified joining tools

UHE Benedikt<fnref n="1" />,ª*, SCHÜTZELT Franziska<fnref n="2" />,ᵇ, DROSSEL Welf-Guntram<fnref n="2" />,ᶜ and MESCHUT Gerson<fnref n="1" />,ᵈ

**Keywords:** Joining, Self-Pierce Riveting, Clinching, Sensitivity Analysis, Unified Tool, Sheet Metal

**Abstract.** Self-pierce riveting and clinching are well-established joining techniques in car body manufacturing. For these two techniques C-frames with drives, which are mounted on industrial robots, are used to realise the joining process. Punch, die and blank holder, which build the joining tools, are characteristic for the mentioned techniques. Even though the joining tools of self-pierce riveting and clinching are similar, joining with the same unified tools is not implemented in car body manufacturing until now. Within this paper, the challenges and requirements of joining with unified joining tools are described. At the beginning of the examination, reference joints are produced experimentally with conventional joining tools. After that, unified joining tools are elaborated, using numerical simulation. At the end, the feasibility of the approach is demonstrated by experimental joining tests with the newly elaborated joining tools.

## Introduction

The mechanical joining technologies *clinching* and *self-pierce riveting* are commonly used joining methods in the sheet metal industry. Both joining methods are suitable for joining two or more joining parts with different thicknesses and tensile strengths, as is used e. g. for joining in composite construction in the automotive industry. There is a recommended joining parts arrangement depending on the joining process. The tool geometry of the punch and the die are different between these two mechanical joining technologies [1].

When changing from clinching to self-pierce riveting, the tool must also be changed. Time and workers are required for the *tool change*, which is problematic for maintaining short cycle times. Apart from this more storage space is required for the two tools, flexibility is decreased and maintenance costs are incurred for both. As the joints are not joined using their preferred joining arrangement, there is a risk of insufficient formation of interlock and neck thickness. In the year 2003, following experimental investigations, Breckweg and Wößner presented a system-side implementation in their patent [2] with a focus on rivet feed and punch drive in order to alternately set clinching and self-pierce riveting joints with a uniform tool.

This paper combines the clinching and self-pierce riveting tools into a *single tool for clinching and self-pierce riveting* without tool change using the same punch-die combination. Optimized tool geometries are derived with the aid of modern numerical simulation methods and sensitivity analysis, similar procedure to the optimization of semi-tubular riveting tools in [3] and the design of tool geometries for clinching, see [4].

<figure><img src="image_1627.png" /></figure>

<footnote n="1">¹ Paderborn University, Laboratory for Material and Joining Technology (LWF), Pohlweg 47-49, 33098 Paderborn, Germany</footnote>

<footnote n="2">² Fraunhofer Institute for Machine Tools and Forming Technology IWU, Nöthnitzer Straße 44, 01187 Dresden, Germany</footnote>

<footnote>a benedikt.uhe@lwf.upb.de, b franziska.schuetzelt@iwu.fraunhofer.de,</footnote>

<footnote>c welf-guntram.drossel@iwu.fraunhofer.de, d gerson.meschut@lwf.upb.de</footnote>
# Flexible self-pierce riveting and clinching with a single joining system using the same unified joining tools

UHE Benedikt<fnref n="1" />,ª*, SCHÜTZELT Franziska<fnref n="2" />,ᵇ, DROSSEL Welf-Guntram<fnref n="2" />,ᶜ and MESCHUT Gerson<fnref n="1" />,ᵈ

**Keywords:** Joining, Self-Pierce Riveting, Clinching, Sensitivity Analysis, Unified Tool, Sheet Metal

**Abstract.** Self-pierce riveting and clinching are well-established joining techniques in car body manufacturing. For these two techniques C-frames with drives, which are mounted on industrial robots, are used to realise the joining process. Punch, die and blank holder, which build the joining tools, are characteristic for the mentioned techniques. Even though the joining tools of self-pierce riveting and clinching are similar, joining with the same unified tools is not implemented in car body manufacturing until now. Within this paper, the challenges and requirements of joining with unified joining tools are described. At the beginning of the examination, reference joints are produced experimentally with conventional joining tools. After that, unified joining tools are elaborated, using numerical simulation. At the end, the feasibility of the approach is demonstrated by experimental joining tests with the newly elaborated joining tools.

## Introduction

The mechanical joining technologies *clinching* and *self-pierce riveting* are commonly used joining methods in the sheet metal industry. Both joining methods are suitable for joining two or more joining parts with different thicknesses and tensile strengths, as is used e. g. for joining in composite construction in the automotive industry. There is a recommended joining parts arrangement depending on the joining process. The tool geometry of the punch and the die are different between these two mechanical joining technologies [1].

When changing from clinching to self-pierce riveting, the tool must also be changed. Time and workers are required for the *tool change*, which is problematic for maintaining short cycle times. Apart from this more storage space is required for the two tools, flexibility is decreased and maintenance costs are incurred for both. As the joints are not joined using their preferred joining arrangement, there is a risk of insufficient formation of interlock and neck thickness. In the year 2003, following experimental investigations, Breckweg and Wößner presented a system-side implementation in their patent [2] with a focus on rivet feed and punch drive in order to alternately set clinching and self-pierce riveting joints with a uniform tool.

This paper combines the clinching and self-pierce riveting tools into a *single tool for clinching and self-pierce riveting* without tool change using the same punch-die combination. Optimized tool geometries are derived with the aid of modern numerical simulation methods and sensitivity analysis, similar procedure to the optimization of semi-tubular riveting tools in [3] and the design of tool geometries for clinching, see [4].

<figure><img src="image_1628.png" /></figure>

<footnote n="1">¹ Paderborn University, Laboratory for Material and Joining Technology (LWF), Pohlweg 47-49, 33098 Paderborn, Germany</footnote>

<footnote n="2">² Fraunhofer Institute for Machine Tools and Forming Technology IWU, Nöthnitzer Straße 44, 01187 Dresden, Germany</footnote>

<footnote>a benedikt.uhe@lwf.upb.de, b franziska.schuetzelt@iwu.fraunhofer.de,</footnote>

<footnote>c welf-guntram.drossel@iwu.fraunhofer.de, d gerson.meschut@lwf.upb.de</footnote>
## Self-pierce Riveting and Clinching

For the self-pierce riveting (SPR), the sheets are clamped between the blank holder and the die. Then the rivet is pressed into the sheets by the punch. After piercing the punch-sided sheet the rivet flares so that an interlock is generated within the die-sided sheet, see Figure 1. At the process end, the punch and the blank holder are run back.

The interlock and the minimum die-sided material thickness are the most important characteristic joint parameters [5]. A sufficient interlock has to be created to ensure an appropriate joint strength. The minimum die-sided material thickness indicates that the die-sided sheet is still intact after joining. This is required to prevent corrosion. Through the closed die side of the joint no media can get into the joint [6]. The geometrical and mechanical properties of the rivet as well as the contour of the die are varied to customise the process configuration to the joining task. The main parameters of the die are the depth and the width. The rivet geometry is mainly characterised by the nominal diameter, the nominal length and the hole contour [7]. With the self-pierce riveting sheet materials up to a tensile strength of 2000 MPa can be joined reliably. The preferred joining arrangement with self-pierce riveting is high strength to low strength and thin to thick [8].

<figure><img src="image_1628.png" /><figcaption>Figure 1: Characteristics of the self-pierce riveting process and the riveted joint</figcaption></figure>

In opposite to the self-pierce riveting, no auxiliary element is required for clinching. To create a joint only a punch and die are needed. As for the self-pierce riveting the sheets are clamped by a blank holder. In the context of this paper, only rigid dies are of interest. The process stages of the clinching process are shown in Figure 2. At the beginning, the sheets are clamped between the die and the blank holder. Then the sheet material is formed into the die by the punch. The material of the die-sided sheet is formed into the ring channel so that the material of the punch-sided sheet can generate an interlock within the die-sided sheet. At the end, the tool is released [9].

The main characteristic joint parameters are the interlock, the neck thickness and the bottom thickness. A minimum bottom thickness has to be ensured for the same reason as for the self-pierce riveting. The interlock and the neck thickness define the joint strength. Through clinching sheet material with a tensile strength up to 600 MPa can be joined. The preferred joining arrangement with clinching is high strength to low strength and thick to thin [10]. The adaptation of the process parameters to the properties of the sheet to be joined can be achieved through adjusting the geometries of the punch and the die. Main parameters to do so are the punch diameter, the contour of the punch, which can include a tip or can be designed conically, and the diameter and the width of the die [11]. For self-pierce riveting as well as for clinching joining systems consisting of a C-frame a drive and a control unit are used [8].
## Self-pierce Riveting and Clinching

For the self-pierce riveting (SPR), the sheets are clamped between the blank holder and the die. Then the rivet is pressed into the sheets by the punch. After piercing the punch-sided sheet the rivet flares so that an interlock is generated within the die-sided sheet, see Figure 1. At the process end, the punch and the blank holder are run back.

The interlock and the minimum die-sided material thickness are the most important characteristic joint parameters [5]. A sufficient interlock has to be created to ensure an appropriate joint strength. The minimum die-sided material thickness indicates that the die-sided sheet is still intact after joining. This is required to prevent corrosion. Through the closed die side of the joint no media can get into the joint [6]. The geometrical and mechanical properties of the rivet as well as the contour of the die are varied to customise the process configuration to the joining task. The main parameters of the die are the depth and the width. The rivet geometry is mainly characterised by the nominal diameter, the nominal length and the hole contour [7]. With the self-pierce riveting sheet materials up to a tensile strength of 2000 MPa can be joined reliably. The preferred joining arrangement with self-pierce riveting is high strength to low strength and thin to thick [8].

<figure><img src="image_1629.png" /><figcaption>Figure 1: Characteristics of the self-pierce riveting process and the riveted joint</figcaption></figure>

In opposite to the self-pierce riveting, no auxiliary element is required for clinching. To create a joint only a punch and die are needed. As for the self-pierce riveting the sheets are clamped by a blank holder. In the context of this paper, only rigid dies are of interest. The process stages of the clinching process are shown in Figure 2. At the beginning, the sheets are clamped between the die and the blank holder. Then the sheet material is formed into the die by the punch. The material of the die-sided sheet is formed into the ring channel so that the material of the punch-sided sheet can generate an interlock within the die-sided sheet. At the end, the tool is released [9].

The main characteristic joint parameters are the interlock, the neck thickness and the bottom thickness. A minimum bottom thickness has to be ensured for the same reason as for the self-pierce riveting. The interlock and the neck thickness define the joint strength. Through clinching sheet material with a tensile strength up to 600 MPa can be joined. The preferred joining arrangement with clinching is high strength to low strength and thick to thin [10]. The adaptation of the process parameters to the properties of the sheet to be joined can be achieved through adjusting the geometries of the punch and the die. Main parameters to do so are the punch diameter, the contour of the punch, which can include a tip or can be designed conically, and the diameter and the width of the die [11]. For self-pierce riveting as well as for clinching joining systems consisting of a C-frame a drive and a control unit are used [8].
<figure><img src="image_1629.png" /><figcaption>Figure 2: Characteristics of the clinching process and the clinched joint</figcaption></figure>

## Selection of Rivets and Clinching Tools

The experimental part of the investigation focusses on a joining task, which comprises the steel CR340LA and the aluminium alloy EN AW-5182. An overview of the properties of the examined sheets is given in Table 1. The yield strength in Table 1 was determined experimentally through compression tests. The selected material combination is joined from two sides, resulting in two cases with different requirements for the joining process and the used joining tools. It has to be recognised that preferred joining directions of self-pierce riveting and clinching are disregarded. Through that, a wide range of possible applications of the tools to be developed can be covered. This is especially relevant for the clinching process, where small changes in the sheet material properties have already a great influence on the interlock as well as on the neck and bottom thicknesses.

<table><caption>Table 1: Properties of the examined sheets</caption><thead><tr><td>Material</td><td>Thickness</td><td>Yield strength</td></tr></thead><tbody><tr><td>CR340LA</td><td>1.5 mm</td><td>Ø 473 MPa*</td></tr><tr><td>EN AW-5182</td><td>2.0 mm</td><td>Ø 108 MPa*</td></tr></tbody></table>

* determined in compression tests
number of experimental tests n = 5

For the selection of suitable rivets and dies for self-pierce riveting on the one hand and appropriate clinching tools on the other hand, experimental joining tests are conducted. To this, a joining system of type TOX® TZ-VSN 08.413026 (TOX® PRESSOTECHNIK GmbH & Co. KG, Germany) is used. The system consists of a C-frame, a servo-electrical drive and a control unit. It allows riveting and clinching with a single joining system by using different tool holders.

Rivets and dies from Tucker® (STANLEY Engineered Fastening Tucker GmbH, Germany) are used for self-pierce riveting. For clinching tools from TOX® (TOX® PRESSOTECHNIK GmbH & Co. KG, Germany) are used. The produced joints with the selected rivets, dies and clinching tools are shown in Table 2. The C-rivet is common for joining tasks with mild steel and aluminium. The selected rivets for the two different kinds of joints only differ in their length. The rivet material 38B2 is quenched and tempered to a hardness of 480 ± 30 HV10, which corresponds to the hardness class H4 from [8].
<figure><img src="image_1630.png" /><figcaption>Figure 2: Characteristics of the clinching process and the clinched joint</figcaption></figure>

## Selection of Rivets and Clinching Tools

The experimental part of the investigation focusses on a joining task, which comprises the steel CR340LA and the aluminium alloy EN AW-5182. An overview of the properties of the examined sheets is given in Table 1. The yield strength in Table 1 was determined experimentally through compression tests. The selected material combination is joined from two sides, resulting in two cases with different requirements for the joining process and the used joining tools. It has to be recognised that preferred joining directions of self-pierce riveting and clinching are disregarded. Through that, a wide range of possible applications of the tools to be developed can be covered. This is especially relevant for the clinching process, where small changes in the sheet material properties have already a great influence on the interlock as well as on the neck and bottom thicknesses.

<table><caption>Table 1: Properties of the examined sheets</caption><thead><tr><td>Material</td><td>Thickness</td><td>Yield strength</td></tr></thead><tbody><tr><td>CR340LA</td><td>1.5 mm</td><td>Ø 473 MPa*</td></tr><tr><td>EN AW-5182</td><td>2.0 mm</td><td>Ø 108 MPa*</td></tr></tbody></table>

* determined in compression tests
number of experimental tests n = 5

For the selection of suitable rivets and dies for self-pierce riveting on the one hand and appropriate clinching tools on the other hand, experimental joining tests are conducted. To this, a joining system of type TOX® TZ-VSN 08.413026 (TOX® PRESSOTECHNIK GmbH & Co. KG, Germany) is used. The system consists of a C-frame, a servo-electrical drive and a control unit. It allows riveting and clinching with a single joining system by using different tool holders.

Rivets and dies from Tucker® (STANLEY Engineered Fastening Tucker GmbH, Germany) are used for self-pierce riveting. For clinching tools from TOX® (TOX® PRESSOTECHNIK GmbH & Co. KG, Germany) are used. The produced joints with the selected rivets, dies and clinching tools are shown in Table 2. The C-rivet is common for joining tasks with mild steel and aluminium. The selected rivets for the two different kinds of joints only differ in their length. The rivet material 38B2 is quenched and tempered to a hardness of 480 ± 30 HV10, which corresponds to the hardness class H4 from [8].
<table><caption>Table 2: Properties of the selected rivets, dies and clinching tools</caption><thead><tr><th rowspan="3"></th><th colspan="4">Joining task</th></tr><tr><th colspan="2">EN AW-5182 + CR340LA</th><th colspan="2">CR340LA + EN AW-5182</th></tr><tr><th>Rivet</th><th>Die</th><th>Rivet</th><th>Die</th></tr></thead><tbody><tr><th>SPR</th><td>Type: C<br>Length: 5.5 mm<br>Material: 38B2 H4<br>Coating: Almac®</td><td>Type: T057<br>Width: 9.5 mm<br>Depth: 2.0 mm</td><td>Type: C<br>Length: 5.0 mm<br>Material: 38B2 H4<br>Coating: Almac®</td><td>Type: T008<br>Width: 9.0 mm<br>Depth: 2.0 mm</td></tr><tr><th>Clinching</th><td>Type: AB60100<br>Diameter: 6.0 mm<br>Shank: cylindrical<br>End face: without tip</td><td>Type: BE8016<br>Width: 8.0 mm<br>Depth: 1.6 mm</td><td>Type: AC56100<br>Diameter: 5.6 mm<br>Shank: conical<br>End face: without tip</td><td>Type: BE8010<br>Width: 8.0 mm<br>Depth: 1.0 mm</td></tr></tbody></table>

The punch for the self-pierce riveting has a diameter of 7.5 mm. The produced joints are shown in the next section in the context of the validation of the numerical models. For the comparison of simulation and experiment, the stiffness of the joining system is eliminated by adjusting the experimentally measured force-displacement-curves. The SPR dies and the clinching tools are used as reference for the development of the same unified tools for both joining techniques. Based on the produced joints, numerical models, which are necessary to elaborate unified joining tool geometries, are created as described in the next section.

## Numerical simulation

The numerical simulation is performed using the commercial FEA software DEFORM v13.0 from SFTC with the integrated Newton-Raphson iteration for implicit calculation. Joining parts and rivet are cross-linked with elastic-plastic material behavior, tools (e. g. punch) are assumed rigid. The numerical simulation is limited to the joining device consisting of the punch, blank holder and die. The elasticity of the joining pliers is not taken into account in the calculation. Analogous to the tools used in the experiment for the two different kinds of joints as shown in the previous chapter, rotationally symmetric 2D models are used as an established method [3, 4] instead of computationally intensive 3D processes. Compared to the edge of the simulated sheet metal finer meshes are generated in areas of high deformation like neck areas and interlock. The geometric separation criterion is used for simulating SPR. The data of the flow curve stored for the simulation was determined experimentally in compression tests. The coefficient of friction was based on own experience from previous parameter studies and data from the literature. The minimal bottom thickness $t_b$ generated during the experiments was used as stopping criterion in the numerical simulation.

The numerical model was validated by comparing it with experimental results, confirming its accurate reproduction of the joint behavior without errors. The green colored contour is the result of the simulation in the following figure. As shown in Figure 3, the load-stroke curves and the characteristic geometrical values were compared between the experiment (gray) and simulation (green) of the two exemplified kinds of joints. The experimental cross-section and calculated contour are almost identical. Small deviations can be observed in the SPR joint of the second joint, particularly within the rivet. It is difficult to achieve a good agreement at this point of the punched-out material. In the simulation, a constant value is assumed as the friction coefficient over the entire sheet metal surface, while the coating inside the rivet may be slightly different, resulting in different behavior due to factors such as different roughness.
<table><caption>Table 2: Properties of the selected rivets, dies and clinching tools</caption><thead><tr><th rowspan="3"></th><th colspan="4">Joining task</th></tr><tr><th colspan="2">EN AW-5182 + CR340LA</th><th colspan="2">CR340LA + EN AW-5182</th></tr><tr><th>Rivet</th><th>Die</th><th>Rivet</th><th>Die</th></tr></thead><tbody><tr><th>SPR</th><td>Type: C<br>Length: 5.5 mm<br>Material: 38B2 H4<br>Coating: Almac®</td><td>Type: T057<br>Width: 9.5 mm<br>Depth: 2.0 mm</td><td>Type: C<br>Length: 5.0 mm<br>Material: 38B2 H4<br>Coating: Almac®</td><td>Type: T008<br>Width: 9.0 mm<br>Depth: 2.0 mm</td></tr><tr><th>Clinching</th><td>Type: AB60100<br>Diameter: 6.0 mm<br>Shank: cylindrical<br>End face: without tip</td><td>Type: BE8016<br>Width: 8.0 mm<br>Depth: 1.6 mm</td><td>Type: AC56100<br>Diameter: 5.6 mm<br>Shank: conical<br>End face: without tip</td><td>Type: BE8010<br>Width: 8.0 mm<br>Depth: 1.0 mm</td></tr></tbody></table>

The punch for the self-pierce riveting has a diameter of 7.5 mm. The produced joints are shown in the next section in the context of the validation of the numerical models. For the comparison of simulation and experiment, the stiffness of the joining system is eliminated by adjusting the experimentally measured force-displacement-curves. The SPR dies and the clinching tools are used as reference for the development of the same unified tools for both joining techniques. Based on the produced joints, numerical models, which are necessary to elaborate unified joining tool geometries, are created as described in the next section.

## Numerical simulation

The numerical simulation is performed using the commercial FEA software DEFORM v13.0 from SFTC with the integrated Newton-Raphson iteration for implicit calculation. Joining parts and rivet are cross-linked with elastic-plastic material behavior, tools (e. g. punch) are assumed rigid. The numerical simulation is limited to the joining device consisting of the punch, blank holder and die. The elasticity of the joining pliers is not taken into account in the calculation. Analogous to the tools used in the experiment for the two different kinds of joints as shown in the previous chapter, rotationally symmetric 2D models are used as an established method [3, 4] instead of computationally intensive 3D processes. Compared to the edge of the simulated sheet metal finer meshes are generated in areas of high deformation like neck areas and interlock. The geometric separation criterion is used for simulating SPR. The data of the flow curve stored for the simulation was determined experimentally in compression tests. The coefficient of friction was based on own experience from previous parameter studies and data from the literature. The minimal bottom thickness $t_b$ generated during the experiments was used as stopping criterion in the numerical simulation.

The numerical model was validated by comparing it with experimental results, confirming its accurate reproduction of the joint behavior without errors. The green colored contour is the result of the simulation in the following figure. As shown in Figure 3, the load-stroke curves and the characteristic geometrical values were compared between the experiment (gray) and simulation (green) of the two exemplified kinds of joints. The experimental cross-section and calculated contour are almost identical. Small deviations can be observed in the SPR joint of the second joint, particularly within the rivet. It is difficult to achieve a good agreement at this point of the punched-out material. In the simulation, a constant value is assumed as the friction coefficient over the entire sheet metal surface, while the coating inside the rivet may be slightly different, resulting in different behavior due to factors such as different roughness.
<table><caption>Figure 3: Comparison of experimental cross-section and simulated contour and resulting characteristic geometrical values of the SPR joint (left) and clinch joint (right) for the two different kinds of joints</caption><thead><tr><th colspan="6">EN AW-5182 t = 2,0 mm in CR340LA t = 1,5 mm</th></tr><tr><th colspan="3">Self-pierce Riveting</th><th colspan="3">Clinching</th></tr><tr><th></th><th>u<sub>1,2</sub> in mm</th><th>t<sub>r</sub> in mm</th><th>F in kN</th><th>Exp.</th><th>f in mm</th><th>t<sub>n</sub> in mm</th><th>F in kN</th></tr></thead><tbody><tr><th>Exp.</th><td>0.32 | 0.35</td><td>0.40</td><td>44.9</td><td>0.22 | 0.26</td><td>0.36 | 0.43</td><td>55.4</td><td>55.4</td></tr><tr><th>Sim.</th><td>0.31</td><td>0.47</td><td>55.0</td><td>0.22</td><td>0.30</td><td>60.0</td><td>60.0</td></tr></tbody></table><table><thead><tr><th colspan="6">CR340LA t = 1,5 mm in EN AW-5182 t = 2,0 mm</th></tr><tr><th colspan="3">Self-pierce Riveting</th><th colspan="3">Clinching</th></tr><tr><th></th><th>u<sub>1,2</sub> in mm</th><th>t<sub>r</sub> in mm</th><th>F in kN</th><th>Exp.</th><th>f in mm</th><th>t<sub>n</sub> in mm</th><th>F in kN</th></tr></thead><tbody><tr><th>Exp.</th><td>0.55 | 0.43</td><td>0.65</td><td>45.3</td><td>0.21 | 0.13</td><td>0.30 | 0.36</td><td>48.8</td><td>48.8</td></tr><tr><th>Sim.</th><td>0.45</td><td>0.64</td><td>50.1</td><td>0.01</td><td>0.40</td><td>52.0</td><td>52.0</td></tr></tbody></table>

The characteristic geometric values generated by the simulation are also almost identical to those of the experiments. This leads to the conclusion that the numerical simulation has high
<table><caption>Figure 3: Comparison of experimental cross-section and simulated contour and resulting characteristic geometrical values of the SPR joint (left) and clinch joint (right) for the two different kinds of joints</caption><thead><tr><th colspan="6">EN AW-5182 t = 2,0 mm in CR340LA t = 1,5 mm</th></tr><tr><th colspan="3">Self-pierce Riveting</th><th colspan="3">Clinching</th></tr><tr><th></th><th>u<sub>1,2</sub> in mm</th><th>t<sub>r</sub> in mm</th><th>F in kN</th><th>f in mm</th><th>t<sub>n</sub> in mm</th><th>F in kN</th></tr></thead><tbody><tr><th>Exp.</th><td>0.32 | 0.35</td><td>0.40</td><td>44.9</td><td>0.22 | 0.26</td><td>0.36 | 0.43</td><td>55.4</td></tr><tr><th>Sim.</th><td>0.31</td><td>0.47</td><td>55.0</td><td>0.22</td><td>0.30</td><td>60.0</td></tr></tbody></table><table><thead><tr><th colspan="6">CR340LA t = 1,5 mm in EN AW-5182 t = 2,0 mm</th></tr><tr><th colspan="3">Self-pierce Riveting</th><th colspan="3">Clinching</th></tr><tr><th></th><th>u<sub>1,2</sub> in mm</th><th>t<sub>r</sub> in mm</th><th>F in kN</th><th>f in mm</th><th>t<sub>n</sub> in mm</th><th>F in kN</th></tr></thead><tbody><tr><th>Exp.</th><td>0.55 | 0.43</td><td>0.65</td><td>45.3</td><td>0.21 | 0.13</td><td>0.30 | 0.36</td><td>48.8</td></tr><tr><th>Sim.</th><td>0.45</td><td>0.64</td><td>50.1</td><td>0.01</td><td>0.40</td><td>52.0</td></tr></tbody></table>

The characteristic geometric values generated by the simulation are also almost identical to those of the experiments. This leads to the conclusion that the numerical simulation has high
predictive power and accuracy and can be used in the following sensitivity analysis instead of costly experiments.

## Sensitivity Analysis

With the help of numerical simulation, a sensitivity analysis is used to evaluate the dependencies between the input and output parameters. The numerical simulation models for the sensitivity analysis are set up parametrically. The Design of Experiments (DoE) of the parameters to be varied is carried out using the commercial statistical software OptiSlang v7.4 from Dynardo. Parameters to be varied are e. g. die diameter from 8 mm to 10 mm, die depth from 1 mm to 2 mm and punch diameter from 5.5 mm to 6.4 mm. For the sensitivity analysis each analyzed joint is tested by 70 numerical simulations. Once the simulation complete, the characteristic geometrical values are measured automatically using the generated joint contour based on a Python script in DEFORM and prepared for the import into the statistical software OptiSlang. This approach has already proven its worth in numerous research projects like [3, 4, 12].

For each investigated joint, a metamodel was computed. Figure 4 shows the resulting metamodel of the sensitivity analysis of clinching joint 1 (EN AW-5182 $t = 2$ mm in CR340LA $t = 1.5$ mm) to show the influence of die depth and die diameter on neck thickness and interlock. Thanks to the combination of multiple metamodels, as done in projects like [4], [12], and [13], an optimal parameter range for the constructive design of the unified tool can be identified for the investigated joints.

<figure><img src="image_1630.png" /><figcaption>Figure 4: Resulting metamodel to show the influence of die parameters on interlock (left) and neck thickness (right)</figcaption></figure>

By reducing the information from the 3D metamodels like Figure 4 to a response surface 2D plot, for example, the influence of die depth and die diameter can be compared across all different joints that have been investigated in this paper, as shown in Figure 5. It can be observed that an increase in die diameter leads to a reduced interlock. In contrast, it can be noted that a deeper die (see blue and gray lines in the plot) results in a greater formation of interlock. Additionally, it can be noted that a die depth of 1.1 mm is not suitable for the investigated joints as it does not result in sufficient interlock formation.

The Coefficient of Prognosis (CoP) of the interlock in the investigated joints is, with approximately 70 % to 90 %, lower than the CoP of the neck thickness, which is around 85 % to
predictive power and accuracy and can be used in the following sensitivity analysis instead of costly experiments.

## Sensitivity Analysis

With the help of numerical simulation, a sensitivity analysis is used to evaluate the dependencies between the input and output parameters. The numerical simulation models for the sensitivity analysis are set up parametrically. The Design of Experiments (DoE) of the parameters to be varied is carried out using the commercial statistical software OptiSlang v7.4 from Dynardo. Parameters to be varied are e. g. die diameter from 8 mm to 10 mm, die depth from 1 mm to 2 mm and punch diameter from 5.5 mm to 6.4 mm. For the sensitivity analysis each analyzed joint is tested by 70 numerical simulations. Once the simulation complete, the characteristic geometrical values are measured automatically using the generated joint contour based on a Python script in DEFORM and prepared for the import into the statistical software OptiSlang. This approach has already proven its worth in numerous research projects like [3, 4, 12].

For each investigated joint, a metamodel was computed. Figure 4 shows the resulting metamodel of the sensitivity analysis of clinching joint 1 (EN AW-5182 $t = 2$ mm in CR340LA $t = 1.5$ mm) to show the influence of die depth and die diameter on neck thickness and interlock. Thanks to the combination of multiple metamodels, as done in projects like [4], [12], and [13], an optimal parameter range for the constructive design of the unified tool can be identified for the investigated joints.

<figure><img src="image_1631.png" /><figcaption>Figure 4: Resulting metamodel to show the influence of die parameters on interlock (left) and neck thickness (right)</figcaption></figure>

By reducing the information from the 3D metamodels like Figure 4 to a response surface 2D plot, for example, the influence of die depth and die diameter can be compared across all different joints that have been investigated in this paper, as shown in Figure 5. It can be observed that an increase in die diameter leads to a reduced interlock. In contrast, it can be noted that a deeper die (see blue and gray lines in the plot) results in a greater formation of interlock. Additionally, it can be noted that a die depth of 1.1 mm is not suitable for the investigated joints as it does not result in sufficient interlock formation.

The Coefficient of Prognosis (CoP) of the interlock in the investigated joints is, with approximately 70 % to 90 %, lower than the CoP of the neck thickness, which is around 85 % to
95 %. Since the interlock is formed relatively late in the joining process, it is correspondingly harder to predict. This should be taken into account when comparing the results of the metamodel analysis and the real experiment. On the other hand, the change in neck thickness occurs continuously with the start of the punch entry into the material and can therefore be predicted better.

<figure><img src="image_1631.png" /><figcaption>Figure 5: Overview of the influence of die depth and die diameter on the formation of interlock</figcaption></figure>

However, for designing the unified tools, it is not only crucial to consider the influence of die parameters on the formation of interlocks. The influence on neck thickness and bottom thickness must also be considered in the tool design. The metamodels demonstrate a mutual influence between interlock and neck thickness, which must be taken into account when selecting the optimal tool parameters for the unified tool. An increase in interlock always leads to a decrease in neck thickness in clinch joints. In addition to the die parameters, the punch parameters like the radius of the punch edge primarily influence the neck thickness of clinch joints. The sensitivity analysis also shows that the choice of punch diameter does not have a negative impact on the formation of the SPR joint, despite having a smaller punch diameter than the rivet diameter.

Furthermore, it is also clearly visible that joint 2 corresponds to the preferred joining direction of SPR. A significant formation of interlocks can be observed in the SPR joint, while the clinch joint shows relatively small interlocks, not exceeding 0.2 mm, despite increasing die depth. On the other hand, it can also be said that even without considering the preferred arrangement of the joining partners, joints can still be generated with optimized tool parameters.
95 %. Since the interlock is formed relatively late in the joining process, it is correspondingly harder to predict. This should be taken into account when comparing the results of the metamodel analysis and the real experiment. On the other hand, the change in neck thickness occurs continuously with the start of the punch entry into the material and can therefore be predicted better.

<figure><img src="image_1632.png" /><figcaption>Figure 5: Overview of the influence of die depth and die diameter on the formation of interlock</figcaption></figure>

However, for designing the unified tools, it is not only crucial to consider the influence of die parameters on the formation of interlocks. The influence on neck thickness and bottom thickness must also be considered in the tool design. The metamodels demonstrate a mutual influence between interlock and neck thickness, which must be taken into account when selecting the optimal tool parameters for the unified tool. An increase in interlock always leads to a decrease in neck thickness in clinch joints. In addition to the die parameters, the punch parameters like the radius of the punch edge primarily influence the neck thickness of clinch joints. The sensitivity analysis also shows that the choice of punch diameter does not have a negative impact on the formation of the SPR joint, despite having a smaller punch diameter than the rivet diameter.

Furthermore, it is also clearly visible that joint 2 corresponds to the preferred joining direction of SPR. A significant formation of interlocks can be observed in the SPR joint, while the clinch joint shows relatively small interlocks, not exceeding 0.2 mm, despite increasing die depth. On the other hand, it can also be said that even without considering the preferred arrangement of the joining partners, joints can still be generated with optimized tool parameters.
## Experimental verification of the approach

After the simulation-based development of the unified joining tools, the numerical results are verified by experimental tests. This is done for the considered joining task as explained above. With both joints and both joining techniques, a clinching die of type BE8014 is used. The diameter of the punch is 6.0 mm. The results of the experimental joining tests are shown in Figure 6. It can be seen that the lower punch diameter provokes a small indentation on the rivet head. As the joining force with the same rivet as before remains, the load in the rivet head is increased due to the lower contact area between punch and rivet.

<table><thead><tr><th colspan="3">Joining task</th></tr><tr><th></th><th>Joint 1: EN AW-5182 + CR340LA</th><th>Joint 2: CR340LA + EN AW-5182</th></tr></thead><tbody><tr><th>SPR</th><td>u =<br>0.32 mm<br>t<sub>r</sub> =<br>0.25 mm</td><td>u =<br>0.33 mm<br>t<sub>r</sub> =<br>0.41 mm</td></tr><tr><th>Clinching</th><td>u =<br>0.33 mm<br>f =<br>0.54 mm<br>t<sub>b1</sub> =<br>0.17 mm<br>t<sub>b2</sub> =<br>0.10 mm</td><td>u =<br>0.16 mm<br>f =<br>0.25 mm<br>t<sub>b1</sub> =<br>0.81 mm<br>t<sub>b2</sub> =<br>0.65 mm</td></tr></tbody></table>

Punch: Newly manufactured with a diameter of 6.0 mm

Die: Conventional clinching die BE8014 with a width of 8.0 mm and a depth of 1.4 mm

<caption>Figure 6: Cross-sections of the joints produced with the unified joining tools</caption>

As the diameter of clinching punches is always lower than the diameter of 7.5 mm of the riveting punch, an optical impairment has to be accepted when unified joining tools are used for self-pierce riveting. As the rivet head always represented an optical impairment of the sheet surface, this effect can be tolerated. An impact on the mechanically properties of the rivet and the joint is not expectable, because of the very low deformation in an area that is not relevant for the joint strength. The clinching dies are convenient for self-pierce riveting as well. The clinched joint does also meet all quality criteria. Thus, it can be stated that both material combinations can be joined by self-pierce riveting and clinching with unified joining tools.

## Summary and Outlook

Within this paper, an approach for the unification of the joining tools for self-pierce riveting and clinching was presented. The challenges and requirements of joining with the same unified tools for both joining techniques were described. The feasibility of the approach was demonstrated for two different joining tasks. The chosen material combination consists of steel and aluminium. First, suitable rivets and joining tools were selected and reference joints were produced experimentally. Then, numerical models were created and unified joining tools were elaborated based on a sensitivity analysis. Finally, the unified joining tools were used for the chosen joining
## Experimental verification of the approach

After the simulation-based development of the unified joining tools, the numerical results are verified by experimental tests. This is done for the considered joining task as explained above. With both joints and both joining techniques, a clinching die of type BE8014 is used. The diameter of the punch is 6.0 mm. The results of the experimental joining tests are shown in Figure 6. It can be seen that the lower punch diameter provokes a small indentation on the rivet head. As the joining force with the same rivet as before remains, the load in the rivet head is increased due to the lower contact area between punch and rivet.

<table><thead><tr><th colspan="3">Joining task</th></tr><tr><th></th><th>Joint 1: EN AW-5182 + CR340LA</th><th>Joint 2: CR340LA + EN AW-5182</th></tr></thead><tbody><tr><th>SPR</th><td>u =<br>0.32 mm<br>t<sub>r</sub> =<br>0.25 mm</td><td>u =<br>0.33 mm<br>t<sub>r</sub> =<br>0.41 mm</td></tr><tr><th>Clinching</th><td>u =<br>0.33 mm<br>f =<br>0.54 mm<br>t<sub>b1</sub> =<br>0.17 mm<br>t<sub>b2</sub> =<br>0.10 mm</td><td>u =<br>0.16 mm<br>f =<br>0.25 mm<br>t<sub>b1</sub> =<br>0.81 mm<br>t<sub>b2</sub> =<br>0.65 mm</td></tr></tbody></table>

Punch: Newly manufactured with a diameter of 6.0 mm

Die: Conventional clinching die BE8014 with a width of 8.0 mm and a depth of 1.4 mm

<caption>Figure 6: Cross-sections of the joints produced with the unified joining tools</caption>

As the diameter of clinching punches is always lower than the diameter of 7.5 mm of the riveting punch, an optical impairment has to be accepted when unified joining tools are used for self-pierce riveting. As the rivet head always represented an optical impairment of the sheet surface, this effect can be tolerated. An impact on the mechanically properties of the rivet and the joint is not expectable, because of the very low deformation in an area that is not relevant for the joint strength. The clinching dies are convenient for self-pierce riveting as well. The clinched joint does also meet all quality criteria. Thus, it can be stated that both material combinations can be joined by self-pierce riveting and clinching with unified joining tools.

## Summary and Outlook

Within this paper, an approach for the unification of the joining tools for self-pierce riveting and clinching was presented. The challenges and requirements of joining with the same unified tools for both joining techniques were described. The feasibility of the approach was demonstrated for two different joining tasks. The chosen material combination consists of steel and aluminium. First, suitable rivets and joining tools were selected and reference joints were produced experimentally. Then, numerical models were created and unified joining tools were elaborated based on a sensitivity analysis. Finally, the unified joining tools were used for the chosen joining
task and it could be shown that that self-pierce riveting and clinching with unified joining tools is feasible.

Based on the results presented in this paper, the authors plan to develop unified joining tools for ten different material combinations in total. When the unified joining tools are manufactured, a comparison of experimentally produced joints will be conducted. After that, a comparison of the joint strength of joints produced with conventional and with unified tools is planned.

## Acknowledgements

The results presented are taken from IGF research project 22047 BG of the European Research Association for Sheet Metal Working (EFB), which has been funded by the German Federation of Industrial Research Associations (AiF) under the program for promotion of industrial research (IGF) run by the Federal Ministry for Economic Affairs and Climate Action (BMWK) based on a decision by the German Bundestag.

## References

[1] G. Meschut, V. Janzen, T. Olfermann: Innovative and Highly Productive Joining Technologies for Multi-Material Lightweight Car Body Structures. Journal of Materials Engineering and Performance, 23 (2014) p. 1515-1523. https://doi.org/10.1007/s11665-014-0962-3

[2] A. Breckweg, J. Wößner: Vorrichtung zum Verbinden von plattenförmigen Bauteilen mit variabler Fügetechnik am 15.07.2003 (DE10332124B4). Deutsches Patent- und Markenamt.

[3] W.-G. Drossel, M. Israel: Sensitivitätsanalyse und Robustheitsbewertung beim mechanischen Fügen. EFB-Forschungsbericht, No. 376, 2013. EFB, Hannover.

[4] T. Falk, T. Kropp, W.-G. Drossel: Clinching with divided punch to prevent critical neck thicknesses. Materials Research Proceedings Series, Volume 28, p. 1059-1066, 2023. 1st ed. Millersville: Materials Research Forum LLC. https://doi.org/10.21741/9781644902479-116

[5] D. Li, I. Chrysanthou, G. Patel, G. Williams: Self-piercing riveting - A review, Int. J. Adv. Manuf. Technol. 92 (2017) p. 1777-1824. https://doi.org/10.1007/s00170-017-0156-x

[6] H. Ang: An Overview of Self-pierce riveting Process with Focus on Joint Failures, Corrosion Issues and Optimisation Techniques, Chin. J. Mech. Eng. 34 (2021) 2. https://doi.org/10.1186/s10033-020-00526-3

[7] K. Mori, Y. Abe: A review on mechanical joining of aluminium and high strength steel sheets by plastic deformation, Int. J. Lightweight Mater. Manuf 1 (2018) p. 1-11. https://doi.org/10.1016/j.ijlmm.2018.02.002

[8] Technical Bulletin DVS 3410, Self-Pierce Riveting - Overview, DVS Media: Düsseldorf, Germany, 2019.

[9] M. Eshtayeh, M. Hrairi, A. Mohiuddin: Clinching process for joining dissimilar materials: state of the art, Int. J. Adv. Manuf. Technol. 82 (2016), p. 179-195. https://doi.org/10.1007/s00170-015-7363-0

[10] DVS-EFB Merkblatt 3420: Clinchen-Überblick, 2021, DVS Media, Düsseldorf.

[11] B. Shi, Y. Wang, S. Liu, H. Tian: Design method of the parameters of tools for clinching technology, 455-456 (2012) p. 1491-1496. https://doi.org/10.4028/scientific5/AMR.455-456.1491

[12] T. Falk, W.-G. Drossel: Methodische Flexibilitätssteigerung beim Halbhohlstanznieten. EFB-Forschungsbericht, No. 513, 2019. EFB, Hannover.

[13] T. Falk, T. Kropp, M. Jäckel: FEM simulation for the development of data-based process models in mechanical joining technology. 10th International Seminar & Conference on Advances in Resistance Welding & Mechanical Joining 2018. Esslingen, Germany, 12-14th September 2018.
task and it could be shown that that self-pierce riveting and clinching with unified joining tools is feasible.

Based on the results presented in this paper, the authors plan to develop unified joining tools for ten different material combinations in total. When the unified joining tools are manufactured, a comparison of experimentally produced joints will be conducted. After that, a comparison of the joint strength of joints produced with conventional and with unified tools is planned.

## Acknowledgements

The results presented are taken from IGF research project 22047 BG of the European Research Association for Sheet Metal Working (EFB), which has been funded by the German Federation of Industrial Research Associations (AiF) under the program for promotion of industrial research (IGF) run by the Federal Ministry for Economic Affairs and Climate Action (BMWK) based on a decision by the German Bundestag.

## References

[1] G. Meschut, V. Janzen, T. Olfermann: Innovative and Highly Productive Joining Technologies for Multi-Material Lightweight Car Body Structures. Journal of Materials Engineering and Performance, 23 (2014) p. 1515-1523. https://doi.org/10.1007/s11665-014-0962-3

[2] A. Breckweg, J. Wößner: Vorrichtung zum Verbinden von plattenförmigen Bauteilen mit variabler Fügetechnik am 15.07.2003 (DE10332124B4). Deutsches Patent- und Markenamt.

[3] W.-G. Drossel, M. Israel: Sensitivitätsanalyse und Robustheitsbewertung beim mechanischen Fügen. EFB-Forschungsbericht, No. 376, 2013. EFB, Hannover.

[4] T. Falk, T. Kropp, W.-G. Drossel: Clinching with divided punch to prevent critical neck thicknesses. Materials Research Proceedings Series, Volume 28, p. 1059-1066, 2023. 1st ed. Millersville: Materials Research Forum LLC. https://doi.org/10.21741/9781644902479-116

[5] D. Li, I. Chrysanthou, G. Patel, G. Williams: Self-piercing riveting - A review, Int. J. Adv. Manuf. Technol. 92 (2017) p. 1777-1824. https://doi.org/10.1007/s00170-017-0156-x

[6] H. Ang: An Overview of Self-pierce riveting Process with Focus on Joint Failures, Corrosion Issues and Optimisation Techniques, Chin. J. Mech. Eng. 34 (2021) 2. https://doi.org/10.1186/s10033-020-00526-3

[7] K. Mori, Y. Abe: A review on mechanical joining of aluminium and high strength steel sheets by plastic deformation, Int. J. Lightweight Mater. Manuf 1 (2018) p. 1-11. https://doi.org/10.1016/j.ijlmm.2018.02.002

[8] Technical Bulletin DVS 3410, Self-Pierce Riveting - Overview, DVS Media: Düsseldorf, Germany, 2019.

[9] M. Eshtayeh, M. Hrairi, A. Mohiuddin: Clinching process for joining dissimilar materials: state of the art, Int. J. Adv. Manuf. Technol. 82 (2016), p. 179-195. https://doi.org/10.1007/s00170-015-7363-0

[10] DVS-EFB Merkblatt 3420: Clinchen-Überblick, 2021, DVS Media, Düsseldorf.

[11] B. Shi, Y. Wang, S. Liu, H. Tian: Design method of the parameters of tools for clinching technology, 455-456 (2012) p. 1491-1496. https://doi.org/10.4028/scientific5/AMR.455-456.1491

[12] T. Falk, W.-G. Drossel: Methodische Flexibilitätssteigerung beim Halbhohlstanznieten. EFB-Forschungsbericht, No. 513, 2019. EFB, Hannover.

[13] T. Falk, T. Kropp, M. Jäckel: FEM simulation for the development of data-based process models in mechanical joining technology. 10th International Seminar & Conference on Advances in Resistance Welding & Mechanical Joining 2018. Esslingen, Germany, 12-14th September 2018.
# Dissimilar welding of aluminum to AlSi-coated steel sheets using refill friction stir spot welding: A feasibility study

SHEN Tong<sup>1,a*</sup>, FU Banglong<sup>1,b</sup>, CHEN Ting<sup>1,c</sup>, SUHUDDIN Uceu<sup>1,d</sup> and KLUSEMANN Benjamin<sup>1,2,e</sup>

<sup>1</sup>Helmholtz-Zentrum Hereon, Institute of Material Mechanics, Solid State Materials Processing, Max-Planck-Str. 1, Geesthacht 21502, Germany

<sup>2</sup> Leuphana University Lüneburg, Institute for Production Technology and Systems, Universitätsallee 1, Lüneburg 21335, Germany

<sup>a</sup>tong.shen@hereon.de, <sup>b</sup>banglong.fu@hereon.de, <sup>c</sup>ting.chen@hereon.de,

<sup>d</sup>uceu.suhuddin@hereon.de, <sup>e</sup>benjamin.klusemann@hereon.de

**Keywords:** Refill Friction Stir Spot Welding, Aluminum, Steel, Dissimilar Materials Joining Lap-Shear Properties

**Abstract.** The main objective of this study was to assess and enhance the weldability of AA6016-T4 to AlSi-coated press-hardened steel (PHS), with a focus on investigating the impact of the AlSi coating on the mechanical properties of welded joints. Variations in tool rotation speed, dwell time, and plunge speed were performed to observe their effects on the joint behavior. It was observed that the protective AlSi-coating significantly influences the mechanical strength of the welded joint. Notably, the removal of the AlSi coating resulted in a substantial increase in joint mechanical strength. To systematically assess joint lap-shear strength under static loading conditions, Design of Experiment technique was employed. The influence of the process parameters on joint strength was thoroughly evaluated.

**Introduction**

In response to the escalating threat of global warming and the imperative to reduce greenhouse gas emissions, the EU's 2030 climate and energy framework targets a substantial 55% reduction in emissions compared to 1990 levels [1]. To get on track with the goals in the 2030 scenario, light-weighting of vehicles has been adopted by both governments and automotive industries as one of the most promising approaches. Over the past two decades, the utilization of modern advanced high-strength steels (AHSS) and press-hardening steel (PHS) have allowed for thinner/lighter body-in-white (BiW) structures [2]. These steel structures can be further combined with aluminum (Al) cast nodes and extrusions to reduce the weight effectively without compromising the costs, safety, and performance.

Although welding techniques for Al/steel dissimilar material joints have been developed to address problems in mechanical fastening and adhesive bonding, the application of the welding processes to join steel to aluminum is still difficult. The thermophysical properties of Al and iron (Fe) [3,4] show great differences in melting/boiling points, thermal expansion coefficient, thermal conductivity, specific heat, etc., as a result, defects as cracks, porosities, distortions occur in conventional fusion welding methods. Most importantly, according to the Al-Fe equilibrium phase diagram [5], Al and Fe have limited mutual solubility, thus hard and brittle intermetallic compounds (IMCs) form easily, which decrease the performance of the joint severely. According to the measured properties of Fe-Al IMCs [6], the Fe-rich IMCs (Fe₃Al, FeAl) show much lower hardness than the Al-rich IMCs (Fe₂Al₇, FeAl₂, Fe₂Al₅, and FeAl₃), which are preferable in Al/steel dissimilar welding. However, based on the thermodynamic calculations [7], Al-rich IMCs, especially the Fe₂Al₅, are more prone to form due to the lower reaction enthalpy. Moreover, the

<figure><img src="image_1632.png" /></figure>
# Dissimilar welding of aluminum to AlSi-coated steel sheets using refill friction stir spot welding: A feasibility study

SHEN Tong<sup>1,a*</sup>, FU Banglong<sup>1,b</sup>, CHEN Ting<sup>1,c</sup>, SUHUDDIN Uceu<sup>1,d</sup> and KLUSEMANN Benjamin<sup>1,2,e</sup>

<sup>1</sup>Helmholtz-Zentrum Hereon, Institute of Material Mechanics, Solid State Materials Processing, Max-Planck-Str. 1, Geesthacht 21502, Germany

<sup>2</sup> Leuphana University Lüneburg, Institute for Production Technology and Systems, Universitätsallee 1, Lüneburg 21335, Germany

<sup>a</sup>tong.shen@hereon.de, <sup>b</sup>banglong.fu@hereon.de, <sup>c</sup>ting.chen@hereon.de,

<sup>d</sup>uceu.suhuddin@hereon.de, <sup>e</sup>benjamin.klusemann@hereon.de

**Keywords:** Refill Friction Stir Spot Welding, Aluminum, Steel, Dissimilar Materials Joining Lap-Shear Properties

**Abstract.** The main objective of this study was to assess and enhance the weldability of AA6016-T4 to AlSi-coated press-hardened steel (PHS), with a focus on investigating the impact of the AlSi coating on the mechanical properties of welded joints. Variations in tool rotation speed, dwell time, and plunge speed were performed to observe their effects on the joint behavior. It was observed that the protective AlSi-coating significantly influences the mechanical strength of the welded joint. Notably, the removal of the AlSi coating resulted in a substantial increase in joint mechanical strength. To systematically assess joint lap-shear strength under static loading conditions, Design of Experiment technique was employed. The influence of the process parameters on joint strength was thoroughly evaluated.

**Introduction**

In response to the escalating threat of global warming and the imperative to reduce greenhouse gas emissions, the EU's 2030 climate and energy framework targets a substantial 55% reduction in emissions compared to 1990 levels [1]. To get on track with the goals in the 2030 scenario, light-weighting of vehicles has been adopted by both governments and automotive industries as one of the most promising approaches. Over the past two decades, the utilization of modern advanced high-strength steels (AHSS) and press-hardening steel (PHS) have allowed for thinner/lighter body-in-white (BiW) structures [2]. These steel structures can be further combined with aluminum (Al) cast nodes and extrusions to reduce the weight effectively without compromising the costs, safety, and performance.

Although welding techniques for Al/steel dissimilar material joints have been developed to address problems in mechanical fastening and adhesive bonding, the application of the welding processes to join steel to aluminum is still difficult. The thermophysical properties of Al and iron (Fe) [3,4] show great differences in melting/boiling points, thermal expansion coefficient, thermal conductivity, specific heat, etc., as a result, defects as cracks, porosities, distortions occur in conventional fusion welding methods. Most importantly, according to the Al-Fe equilibrium phase diagram [5], Al and Fe have limited mutual solubility, thus hard and brittle intermetallic compounds (IMCs) form easily, which decrease the performance of the joint severely. According to the measured properties of Fe-Al IMCs [6], the Fe-rich IMCs (Fe₃Al, FeAl) show much lower hardness than the Al-rich IMCs (Fe₂Al₇, FeAl₂, Fe₂Al₅, and FeAl₃), which are preferable in Al/steel dissimilar welding. However, based on the thermodynamic calculations [7], Al-rich IMCs, especially the Fe₂Al₅, are more prone to form due to the lower reaction enthalpy. Moreover, the

<figure><img src="image_1633.png" /></figure>
crystal structure of Fe₂Al₅ is characterized with 30% vacancies along the c-axis [8], it grows rapidly and irregularly, which will decrease the performance of the joint. Several attempts have been conducted to control the heat input in fusion welding methods to reduce the formation of Al-rich IMCs and then improve the strength, which include cold metal transfer [9], implementation of high-power beams like electron beam [10] and laser beam [11], as well as laser-arc hybrid source [12]. The brazing-fusion welding method, i.e. the Al and filler metals (if used) melt while the steel is kept in solid state, can regulate the IMCs effectively. For instance, in laser brazing-fusion welding with Al-Si-Zn filler wire [13], the thickness of Fe-Al IMCs was suppressed below 10 µm due to the absence of melting of steel, and at the same time, the brittleness of the IMCs was decreased with the formation of Fe-Al-Si-Zn IMCs, leading to an improved mechanical strength of the weld.

As a solid-state welding process, friction stir welding (FSW) can eliminate the common problems in fusion welding due to the absence of melting and solidification. Additionally, a thin IMC layer can be produced with low heat input during FSW of Al/Ti [14]. These advantages have benefited its successful industrial production for Al/steel dissimilar components [15], where, compared to conventional welding, FSWed joints showed increased strength, absence of distortion, elimination of environmental pollution, higher space saving with reduced energy consumption by 50 %. Currently, it is estimated that around 85 % of the welding work in the automotive industry is related to spot welds [16]. Thus, the modified spot welding method based on FSW, friction stir spot welding (FSSW), was developed. Although Al/steel spot welds can be produced by FSSW [17], an exit hole is left on the surface of the joint, resulting in problems such as decreased mechanical properties, potential stress concentration preferred sites for corrosion, etc.

As a solution, the refill friction stir spot welding (refill FSSW), a solid-state spot welding process to produce joints without an exit hole, was developed [18]. Refill FSSW is characterized by high energy efficiency and does not require additional filler materials, shielding gases for the operation, etc. The refill FSSW process has two variants depending on the plunge component into the material: probe plunge (PP) and shoulder plunge (SP) variant. The shoulder has a larger tip and side surfaces than the probe, so the SP variant can introduce a higher heat input and produce a larger joint than the PP variant. This results in more efficient material stirring and leads to a stronger joint. The SP variant is used in this study, it consists of four stages, as schematically illustrated in Fig. 1.

<figure><img src="image_1633.png" /><figcaption>Fig. 1 Schematic illustration of SP variant of refill FSSW process</figcaption></figure>

When applying refill FSSW to Al/press-hardened steel (PHS) with AlSi-coating, Ding et al. [19] analyzed the failure behavior and found that, the cracks mainly propagated through the AlSi-coating and AlSi/Al-Fe-Si interface. To enhance the understanding of the impact of AlSi coating on the weldability of Al/PHS steel joining, the current study conducts a direct comparison of the mechanical behavior of joints with varying PHS surface conditions. Furthermore, this study also
crystal structure of Fe₂Al₅ is characterized with 30% vacancies along the c-axis [8], it grows rapidly and irregularly, which will decrease the performance of the joint. Several attempts have been conducted to control the heat input in fusion welding methods to reduce the formation of Al-rich IMCs and then improve the strength, which include cold metal transfer [9], implementation of high-power beams like electron beam [10] and laser beam [11], as well as laser-arc hybrid source [12]. The brazing-fusion welding method, i.e. the Al and filler metals (if used) melt while the steel is kept in solid state, can regulate the IMCs effectively. For instance, in laser brazing-fusion welding with Al-Si-Zn filler wire [13], the thickness of Fe-Al IMCs was suppressed below 10 µm due to the absence of melting of steel, and at the same time, the brittleness of the IMCs was decreased with the formation of Fe-Al-Si-Zn IMCs, leading to an improved mechanical strength of the weld.

As a solid-state welding process, friction stir welding (FSW) can eliminate the common problems in fusion welding due to the absence of melting and solidification. Additionally, a thin IMC layer can be produced with low heat input during FSW of Al/Ti [14]. These advantages have benefited its successful industrial production for Al/steel dissimilar components [15], where, compared to conventional welding, FSWed joints showed increased strength, absence of distortion, elimination of environmental pollution, higher space saving with reduced energy consumption by 50 %. Currently, it is estimated that around 85 % of the welding work in the automotive industry is related to spot welds [16]. Thus, the modified spot welding method based on FSW, friction stir spot welding (FSSW), was developed. Although Al/steel spot welds can be produced by FSSW [17], an exit hole is left on the surface of the joint, resulting in problems such as decreased mechanical properties, potential stress concentration preferred sites for corrosion, etc.

As a solution, the refill friction stir spot welding (refill FSSW), a solid-state spot welding process to produce joints without an exit hole, was developed [18]. Refill FSSW is characterized by high energy efficiency and does not require additional filler materials, shielding gases for the operation, etc. The refill FSSW process has two variants depending on the plunge component into the material: probe plunge (PP) and shoulder plunge (SP) variant. The shoulder has a larger tip and side surfaces than the probe, so the SP variant can introduce a higher heat input and produce a larger joint than the PP variant. This results in more efficient material stirring and leads to a stronger joint. The SP variant is used in this study, it consists of four stages, as schematically illustrated in Fig. 1.

<figure><img src="image_1634.png" /><figcaption>Fig. 1 Schematic illustration of SP variant of refill FSSW process</figcaption></figure>

When applying refill FSSW to Al/press-hardened steel (PHS) with AlSi-coating, Ding et al. [19] analyzed the failure behavior and found that, the cracks mainly propagated through the AlSi-coating and AlSi/Al-Fe-Si interface. To enhance the understanding of the impact of AlSi coating on the weldability of Al/PHS steel joining, the current study conducts a direct comparison of the mechanical behavior of joints with varying PHS surface conditions. Furthermore, this study also
presents the significance of refill FSSW parameters, namely rotational speed (RotSpeed), dwell time (DweTime), and plunge speed (PluSpeed), in relation to the weld strength of the resulting joint.

## Material and Methods

The AlSi-coated steel sheets utilized in this study were AlSi-coated PHS1500 steel, having a thickness of 1.5 mm and an average AlSi coating mass of 80 g/m². The chemical composition is listed in Tab. 1. To compare the weld outcomes with and without the presence of AlSi coating on the steel, a diamond disk grinder is utilized to eliminate the coating layer. The aluminum plates used in this study were AA6016-T4 aluminum alloy with a thickness of 1.2 mm. Overlap single joints were carried out using a Harms & Wende® RPS100-v32 machine, employing a non-consumable welding tool with probe, shoulder, and clamping ring diameters of 6 mm, 9 mm, and 17 mm, respectively. The employed process parameters are given in the following section.

<table><caption>Tab. 1 Chemical composition of AlSi-coated PHS1500 steel (wt %)</caption><thead><tr><th>Steel Grade</th><th>C</th><th>Si</th><th>Mn</th><th>P</th><th>S</th><th>Al</th><th>Cu</th><th>Cr</th><th>N</th><th>Ti+Nb</th><th>Cr+Mo</th><th>B</th></tr></thead><tbody><tr><td>PHS1500</td><td>0.19 - 0.25</td><td>1.00 - 1.70</td><td>0.030</td><td>0.005</td><td>0.15 - 0.50</td><td>0.35</td><td>0.35</td><td>0.02 - 0.05</td><td>0.02 - 0.06</td><td>0.009</td><td>0.19 - 0.25</td><td>1.00 - 1.70</td></tr></tbody></table>

Standard metallographic procedures were applied using Struers Tegramin-30 equipment to investigate the polished cross-sections of the obtained joints. The macro- and microstructures of the joints were observed via a Keyence VHX-6000 optical microscope (OM) and a FEI Quanta 650 FEG scanning electron microscope (SEM), which was fitted with an energy-dispersive X-ray spectroscopy (EDS) and electron backscatter diffraction (EBSD) detector. Quasi-static lap shear tests were conducted using a Zwick/Roell 1478 universal testing machine with a consistent cross-head speed of 1 mm/min at ambient temperature. All testing samples were fabricated into coupons measuring 90 mm x 115 mm, according to the DIN EN ISO 14273 standard [20]. Design of Experiment is employed as a robust statistical tool commonly employed in the analysis and optimization of industrial processes. With the support of the statistical analysis software Minitab, the number of experiments necessary for each design method can be generated. It also enables the analysis of how the performance of the weld responds to variations in welding variables. In this work, a three-factor three-level Box-Behnken experimental Design (BBD) has been employed.

## Results and Discussion

The PHS sheets employed in this study are coated with a consistent layer of AlSi, maintaining a thickness of 30 ± 2 µm. An ultra-thin layer of Fe-Al inter-metallic compound (IMC), measuring 9 ± 1 µm, is present between the AlSi coating and the steel matrix, as illustrated in Fig. 3(a). Fig. 4 illustrates the EDS concentration profiles across the coating layer to the steel substrate. The line scanning result revealed the presence of multiple sublayers distinguished by their chemical compositions and thicknesses. To assess the weldability of AlSi-coated PHS, the impact of different surface conditions on the mechanical properties of the joints is investigated. Two distinct surface conditions of PHS sheets were produced: 1) The original PHS sheet with the AlSi coating, without any additional surface treatment. 2) Mechanically grinding all the coating layers, exposing the PHS steel matrix, as shown in Fig. 3(b).
presents the significance of refill FSSW parameters, namely rotational speed (RotSpeed), dwell time (DweTime), and plunge speed (PluSpeed), in relation to the weld strength of the resulting joint.

## Material and Methods

The AlSi-coated steel sheets utilized in this study were AlSi-coated PHS1500 steel, having a thickness of 1.5 mm and an average AlSi coating mass of 80 g/m². The chemical composition is listed in Tab. 1. To compare the weld outcomes with and without the presence of AlSi coating on the steel, a diamond disk grinder is utilized to eliminate the coating layer. The aluminum plates used in this study were AA6016-T4 aluminum alloy with a thickness of 1.2 mm. Overlap single joints were carried out using a Harms & Wende® RPS100-v32 machine, employing a non-consumable welding tool with probe, shoulder, and clamping ring diameters of 6 mm, 9 mm, and 17 mm, respectively. The employed process parameters are given in the following section.

<table><caption>Tab. 1 Chemical composition of AlSi-coated PHS1500 steel (wt %)</caption><thead><tr><th>Steel Grade</th><th>C</th><th>Si</th><th>Mn</th><th>P</th><th>S</th><th>Al</th><th>Cu</th><th>Cr</th><th>N</th><th>Ti+Nb</th><th>Cr+Mo</th><th>B</th></tr></thead><tbody><tr><td>PHS1500</td><td>0.19 - 0.25</td><td>1.00 - 1.70</td><td>0.030</td><td>0.005</td><td>0.15 - 0.50</td><td>0.35</td><td>0.35</td><td>0.02 - 0.05</td><td>0.02 - 0.06</td><td>0.009</td><td>0.19 - 0.25</td><td>1.00 - 1.70</td></tr></tbody></table>

Standard metallographic procedures were applied using Struers Tegramin-30 equipment to investigate the polished cross-sections of the obtained joints. The macro- and microstructures of the joints were observed via a Keyence VHX-6000 optical microscope (OM) and a FEI Quanta 650 FEG scanning electron microscope (SEM), which was fitted with an energy-dispersive X-ray spectroscopy (EDS) and electron backscatter diffraction (EBSD) detector. Quasi-static lap shear tests were conducted using a Zwick/Roell 1478 universal testing machine with a consistent cross-head speed of 1 mm/min at ambient temperature. All testing samples were fabricated into coupons measuring 90 mm x 115 mm, according to the DIN EN ISO 14273 standard [20]. Design of Experiment is employed as a robust statistical tool commonly employed in the analysis and optimization of industrial processes. With the support of the statistical analysis software Minitab, the number of experiments necessary for each design method can be generated. It also enables the analysis of how the performance of the weld responds to variations in welding variables. In this work, a three-factor three-level Box-Behnken experimental Design (BBD) has been employed.

## Results and Discussion

The PHS sheets employed in this study are coated with a consistent layer of AlSi, maintaining a thickness of 30 ± 2 µm. An ultra-thin layer of Fe-Al inter-metallic compound (IMC), measuring 9 ± 1 µm, is present between the AlSi coating and the steel matrix, as illustrated in Fig. 3(a). Fig. 4 illustrates the EDS concentration profiles across the coating layer to the steel substrate. The line scanning result revealed the presence of multiple sublayers distinguished by their chemical compositions and thicknesses. To assess the weldability of AlSi-coated PHS, the impact of different surface conditions on the mechanical properties of the joints is investigated. Two distinct surface conditions of PHS sheets were produced: 1) The original PHS sheet with the AlSi coating, without any additional surface treatment. 2) Mechanically grinding all the coating layers, exposing the PHS steel matrix, as shown in Fig. 3(b).
<figure><img src="image_1634.png" /><figcaption>Fig. 3(a). Microstructure of a) AlSi-coated PHS steel (b) PHS steel after AlSi-coating removal</figcaption></figure>

<figure><img src="image_1635.png" /><figcaption>Fig. 4. EDS scanning results of AlSi-coating (a) coating layer of the as-received base material, (b) line scanning result shown in (a)</figcaption></figure>

The AlSi-coated PHS sheets with varied surface conditions were then joined with AA6016-T4 sheets, using refill FSSW. A summary of the employed process parameters is presented in Tab. 2. The peeling off, i.e. absence of bonding, of welded joints has been observed and marked for the corresponding joint in Tab. 2. Six different parameter settings were employed for each surface condition to vary the heat inputs. For each surface condition, six samples were allocated for microstructure analysis, and an additional six samples were designated for weld strength measurements. The lap-shear strength (LSS) comparison for different surface conditions of PHS1500 is plotted in Fig. 5.
<figure><img src="image_1635.png" /><figcaption>Fig. 3(a). Microstructure of a) AlSi-coated PHS steel (b) PHS steel after AlSi-coating removal</figcaption></figure>

<figure><img src="image_1636.png" /><figcaption>Fig. 4. EDS scanning results of AlSi-coating (a) coating layer of the as-received base material, (b) line scanning result shown in (a)</figcaption></figure>

The AlSi-coated PHS sheets with varied surface conditions were then joined with AA6016-T4 sheets, using refill FSSW. A summary of the employed process parameters is presented in Tab. 2. The peeling off, i.e. absence of bonding, of welded joints has been observed and marked for the corresponding joint in Tab. 2. Six different parameter settings were employed for each surface condition to vary the heat inputs. For each surface condition, six samples were allocated for microstructure analysis, and an additional six samples were designated for weld strength measurements. The lap-shear strength (LSS) comparison for different surface conditions of PHS1500 is plotted in Fig. 5.
<table><caption>Tab. 2. Summarization of comparison study of different surface conditions of PHS sheets lap-shear strength of dissimilar AA6016-PHS1500 joints.</caption><thead><tr><th>Weld No.</th><th>RotSpeed (rpm)</th><th>PluSpeed (mm/s)</th><th>DweTime (s)</th><th>LSS (N)</th><th>Surface condition</th></tr></thead><tbody><tr><td>GT01.og</td><td>2400</td><td>1</td><td>0</td><td>1270</td><td rowspan="5">AlSi-coated PHS</td></tr><tr><td>GT02.og</td><td>2400</td><td>1</td><td>1</td><td>558</td></tr><tr><td>GT03.og</td><td>2500</td><td>1</td><td>2</td><td>428</td></tr><tr><td>GT04.og</td><td>2000</td><td>1</td><td>0</td><td>1370*</td></tr><tr><td>GT05.og</td><td>1800</td><td>1</td><td>2</td><td>981</td></tr><tr><td>GT06.og</td><td>1500</td><td>1</td><td>0</td><td>**</td><td rowspan="5">Uncoated PHS</td></tr><tr><td>GT01.C</td><td>2400</td><td>1</td><td>0</td><td>4650</td></tr><tr><td>GT02.C</td><td>2400</td><td>1</td><td>1</td><td>5010</td></tr><tr><td>GT03.C</td><td>2500</td><td>1</td><td>2</td><td>5190</td></tr><tr><td>GT04.C</td><td>2000</td><td>1</td><td>0</td><td>4430</td></tr><tr><td>GT05.C</td><td>1800</td><td>1</td><td>2</td><td>4920</td><td rowspan="2">Uncoated PHS</td></tr><tr><td>GT06.C</td><td>1500</td><td>1</td><td>0</td><td>5500</td></tr></tbody><tfoot><tr><td colspan="6">*Indicates the absence of bonding in the sample.</td></tr></tfoot></table>

<figure><img src="image_1636.png" /><figcaption>Fig. 5 Comparison of lap-shear strength for different surface conditions PHS1500 when joined with AA6016-T4</figcaption></figure>

Results indicate that with the original AlSi coating, weak bonding or even unbonded areas are commonly observed. Upon removing the AlSi coating, the weld strength noticeably increases. The elimination of the coating layer results in a strength improvement of at least threefold. Cross-sectional macrostructures in Fig. 6 depict the typical appearance of the joining interface in the stirred zone (SZ) of samples with and without the coating layer. The macrograph depicting the joint without AlSi coating, as illustrated in Fig. 6(a), reveals an absence of welding defects. However, in the joint with AlSi coating, as depicted in Fig. 6(b), cracks are evident in the weld. Figure 6(a) further highlights the presence of cracks in the shoulder area, suggesting the
<table><caption>Tab. 2. Summarization of comparison study of different surface conditions of PHS sheets lap-shear strength of dissimilar AA6016-PHS1500 joints.</caption><thead><tr><th>Weld No.</th><th>RotSpeed (rpm)</th><th>PluSpeed (mm/s)</th><th>DweTime (s)</th><th>LSS (N)</th><th>Surface condition</th></tr></thead><tbody><tr><td>GT01.og</td><td>2400</td><td>1</td><td>0</td><td>1270</td><td rowspan="6">AlSi-coated PHS</td></tr><tr><td>GT02.og</td><td>2400</td><td>1</td><td>1</td><td>558</td></tr><tr><td>GT03.og</td><td>2500</td><td>1</td><td>2</td><td>428</td></tr><tr><td>GT04.og</td><td>2000</td><td>1</td><td>0</td><td>1370*</td></tr><tr><td>GT05.og</td><td>1800</td><td>1</td><td>2</td><td>981</td></tr><tr><td>GT06.og</td><td>1500</td><td>1</td><td>0</td><td>**</td></tr><tr><td>GT01.C</td><td>2400</td><td>1</td><td>0</td><td>4650</td><td rowspan="6">Uncoated PHS</td></tr><tr><td>GT02.C</td><td>2400</td><td>1</td><td>1</td><td>5010</td></tr><tr><td>GT03.C</td><td>2500</td><td>1</td><td>2</td><td>5190</td></tr><tr><td>GT04.C</td><td>2000</td><td>1</td><td>0</td><td>4430</td></tr><tr><td>GT05.C</td><td>1800</td><td>1</td><td>2</td><td>4920</td></tr><tr><td>GT06.C</td><td>1500</td><td>1</td><td>0</td><td>5500</td></tr></tbody><tfoot><tr><td colspan="6">*Indicates the absence of bonding in the sample.</td></tr></tfoot></table>

<figure><img src="image_1637.png" /><figcaption>Fig. 5 Comparison of lap-shear strength for different surface conditions PHS1500 when joined with AA6016-T4</figcaption></figure>

Results indicate that with the original AlSi coating, weak bonding or even unbonded areas are commonly observed. Upon removing the AlSi coating, the weld strength noticeably increases. The elimination of the coating layer results in a strength improvement of at least threefold. Cross-sectional macrostructures in Fig. 6 depict the typical appearance of the joining interface in the stirred zone (SZ) of samples with and without the coating layer. The macrograph depicting the joint without AlSi coating, as illustrated in Fig. 6(a), reveals an absence of welding defects. However, in the joint with AlSi coating, as depicted in Fig. 6(b), cracks are evident in the weld. Figure 6(a) further highlights the presence of cracks in the shoulder area, suggesting the
challenging establishment of metallurgical bonding between AlSi coating and Al. This hindrance in achieving metallurgical bonding between dissimilar metals during welding can accelerate the rupture of the welded joint under load.

<figure><img src="image_1637.png" /><figcaption>Fig. 6 Typical cross-sectional macrostructures of the joining interface in the stirred zone (SZ) of samples (a) with and (b) without the AlSi coating layer (RotSpeed: 2400 rpm, PluSpeed: 1 mm/s, DweTime 0 s, PluDepth 1.1 mm).</figcaption></figure>

According to the defined parameter windows for dissimilar welding of AA6016 to PHS1500, a BBD was used for parameter optimization. The investigated parameters and the related levels in the BBD are summarized in Tab. 3. The removal of the AlSi coating from the PHS sheets is regarded as surface pre-treatment in the following.

<table><caption>Tab. 3 Investigated refill FSSW parameters and the related levels in the BBD for dissimilar welding of AA6016 to PHS1500. The AlSi coating has been removed from the PHS sheets.</caption><thead><tr><th>Parameters</th><th>-1</th><th>0</th><th>1</th></tr></thead><tbody><tr><td>Rotation speed (rpm)</td><td>1500</td><td>2000</td><td>2500</td></tr><tr><td>Dwell time (s)</td><td>0</td><td>1</td><td>2</td></tr><tr><td>Plunge speed (mm/s)</td><td>0.5</td><td>1</td><td>1.5</td></tr></tbody></table>

The normal probability plot, Fig. 7(a), confirms that the achieved LSS data fulfills the normality assumption. The Pareto chart obtained is shown in Fig. 7(b). According to the Pareto chart, the terms are arranged in descending order based on the decrease of their standardized effect on the mean LSS value. Rotational speed and dwell time are the only statistically significant parameters for the model. It is noteworthy that one of the main parameters, plunge speed, exhibits the least significance among all the terms in the model.
challenging establishment of metallurgical bonding between AlSi coating and Al. This hindrance in achieving metallurgical bonding between dissimilar metals during welding can accelerate the rupture of the welded joint under load.

<figure><img src="image_1638.png" /><figcaption>Fig. 6 Typical cross-sectional macrostructures of the joining interface in the stirred zone (SZ) of samples (a) with and (b) without the AlSi coating layer (RotSpeed: 2400 rpm, PluSpeed: 1 mm/s, DweTime 0 s, PluDepth 1.1 mm).</figcaption></figure>

According to the defined parameter windows for dissimilar welding of AA6016 to PHS1500, a BBD was used for parameter optimization. The investigated parameters and the related levels in the BBD are summarized in Tab. 3. The removal of the AlSi coating from the PHS sheets is regarded as surface pre-treatment in the following.

<table><caption>Tab. 3 Investigated refill FSSW parameters and the related levels in the BBD for dissimilar welding of AA6016 to PHS1500. The AlSi coating has been removed from the PHS sheets.</caption><thead><tr><th>Parameters</th><th>-1</th><th>0</th><th>1</th></tr></thead><tbody><tr><td>Rotation speed (rpm)</td><td>1500</td><td>2000</td><td>2500</td></tr><tr><td>Dwell time (s)</td><td>0</td><td>1</td><td>2</td></tr><tr><td>Plunge speed (mm/s)</td><td>0.5</td><td>1</td><td>1.5</td></tr></tbody></table>

The normal probability plot, Fig. 7(a), confirms that the achieved LSS data fulfills the normality assumption. The Pareto chart obtained is shown in Fig. 7(b). According to the Pareto chart, the terms are arranged in descending order based on the decrease of their standardized effect on the mean LSS value. Rotational speed and dwell time are the only statistically significant parameters for the model. It is noteworthy that one of the main parameters, plunge speed, exhibits the least significance among all the terms in the model.
<figure><img src="image_1638.png" /><figcaption>Fig. 7 (a) Normal probability plot of standardized residuals and (b) Pareto chart analysis of the main effects and interactions for LSS_Mean</figcaption></figure>

The developed regression model was used to analyze the effects of welding parameters. Fig. 8(a) shows the main effects plot for the mean LSS. The effect of the most significant parameter rotational speed is confirmed by substantial variations in the mean value of LSS. The LSS decreases with the increase of rotational speed from 1500 to 2200 rpm and then exhibits an upward trend. Fig. 8(b) shows the interaction plot of combination of parameters on the LSS. The combination of rotational speed and dwell time, acknowledged as the most significant two-way interaction term, shows a strong influence on the LSS. When the rotational speed is higher than 2000 rpm, an increased dwell time is recommended to enhance the weld strength.

<figure><img src="image_1639.png" /><figcaption>Fig. 8 (a) Main effect plots of the parameters on LSS, (b) Interaction plot of the parameters on LSS.</figcaption></figure>

## Summary

This preliminary study has effectively demonstrated the possibility to weld AA6016-T4 to AlSi-coated PHS1500 via refill FSSW. The impact of welding process parameters on the weld lap-shear strength of dissimilar joints was evaluated via statistical analysis. Based on the experimental results, the following conclusions can be drawn:

* The AlSi-coated PHS steel in its initial state exhibits a predominantly weak bond, with an average lap-shear strength of 0.9 kN.

* The elimination of the AlSi coating layer results in higher weld strength, averaging 4.95 kN. Significantly, each sample exhibits at least a threefold increase in strength compared to AlSi-coated surface conditions under identical welding conditions.
<figure><img src="image_1639.png" /><figcaption>Fig. 7 (a) Normal probability plot of standardized residuals and (b) Pareto chart analysis of the main effects and interactions for LSS_Mean</figcaption></figure>

The developed regression model was used to analyze the effects of welding parameters. Fig. 8(a) shows the main effects plot for the mean LSS. The effect of the most significant parameter rotational speed is confirmed by substantial variations in the mean value of LSS. The LSS decreases with the increase of rotational speed from 1500 to 2200 rpm and then exhibits an upward trend. Fig. 8(b) shows the interaction plot of combination of parameters on the LSS. The combination of rotational speed and dwell time, acknowledged as the most significant two-way interaction term, shows a strong influence on the LSS. When the rotational speed is higher than 2000 rpm, an increased dwell time is recommended to enhance the weld strength.

<figure><img src="image_1640.png" /><figcaption>Fig. 8 (a) Main effect plots of the parameters on LSS, (b) Interaction plot of the parameters on LSS.</figcaption></figure>

## Summary

This preliminary study has effectively demonstrated the possibility to weld AA6016-T4 to AlSi-coated PHS1500 via refill FSSW. The impact of welding process parameters on the weld lap-shear strength of dissimilar joints was evaluated via statistical analysis. Based on the experimental results, the following conclusions can be drawn:

* The AlSi-coated PHS steel in its initial state exhibits a predominantly weak bond, with an average lap-shear strength of 0.9 kN.

* The elimination of the AlSi coating layer results in higher weld strength, averaging 4.95 kN. Significantly, each sample exhibits at least a threefold increase in strength compared to AlSi-coated surface conditions under identical welding conditions.
* Within the specified range of welding parameters, rotational speed proved to be the most influential parameter affecting the lap-shear strength of the joints, with the dwell time following as the next significant factor. Conversely, plunge speed exhibited no significant impact on the performance of the joints.

## Acknowledgment

This result is part of the project titled "Friction stir welding for improving joinability of high-performance steels for automotive components to boost green road mobility — Stir4Steel". This project has received funding from the Research Fund for Coal and Steel under grant agreement No 101034068.

## References

[1] COMMUNICATION FROM THE COMMISSION TO THE EUROPEAN PARLIAMENT. A Clean Planet for all A European strategic long-term vision for a prosperous, modern, competitive and climate neutral economy, 2018. https://eur-lex.europa.eu/legal-content/EN/TXT/?uri=CELEX:52018DC0773. Lesch, N. Kwiaton, F.B. Klose, Advanced High Strength Steels (AHSS) for Automotive Applications – Tailored Properties by Smart Microstructural Adjustments, Steel Res. Int. 88 (2017) 1700210. https://doi.org/10.1002/srin.201700210

[2] C. Lesch, N. Kwiaton, F.B. Klose, Advanced High Strength Steels (AHSS) for Automotive Applications – Tailored Properties by Smart Microstructural Adjustments, Steel Res. Int. 88 (2017) 1700210. https://doi.org/10.1002/srin.201700210

[3] A.S.M. International, ed., ASM Handbook Volume 1: Properties and Selection: Irons, Steels, and High-Performance Alloys, 10th edition, ASM International, Materials Park, Ohio, 1990.

[4] A.I.H. Committee, ASM Handbook Volume 2: Properties and Selection: Nonferrous Alloys and Special-purpose Materials, ASM International, 1990. https://books.google.de/books?id=wxA7AQAAIAAJ

[5] M.E. Schlesinger, E.M. Mueller, ASM Handbook, Volume 3: Alloy Phase Diagrams, ASM International, Materials Park, Ohio, 1983.

[6] L. Tricarico, R. Spina, D. Sorgente, et al., Effects of heat treatments on mechanical properties of Fe/Al explosion-welded structural transition joints, Mater. Des. 30 (2009) 2693-2700. https://doi.org/10.1016/j.matdes.2008.10.010

[7] R.W. Richards, R.D. Jones, P.D. Clements, et al., Metallurgy of continuous hot dip aluminizing, Int. Mater. Rev. 39 (1994) 191-212. https://doi.org/10.1179/imr.1994.39.5.191.

[8] J. Yang, J.P. Oliveira, Y. Li, et al., Laser techniques for dissimilar joining of aluminum alloys to steels: A critical review, J. Mater. Process. Technol. 301 (2022). https://doi.org/10.1016/j.jmatprotec.2021.117443

[9] J. Chen, B. Shalchi Amirkhiz, R. Zhang, et al., On the Joint Formation and Interfacial Microstructure of Cold Metal Transfer Cycle Step Braze Welding of Aluminum to Steel Butt Joint, Metall. Mater. Trans. A. 51 (2020) 5198-5212. https://doi.org/10.1007/s11661-020-05917-8

[10] G. Chen, J. Liu, X. Shu, et al., Beam scanning effect on properties optimization of thick-plate 2A12 aluminum alloy electron-beam welding joints, Mater. Sci. Eng. A. 744 (2019) 583-592. https://doi.org/10.1016/j.msea.2018.12.034
* Within the specified range of welding parameters, rotational speed proved to be the most influential parameter affecting the lap-shear strength of the joints, with the dwell time following as the next significant factor. Conversely, plunge speed exhibited no significant impact on the performance of the joints.

## Acknowledgment

This result is part of the project titled "Friction stir welding for improving joinability of high-performance steels for automotive components to boost green road mobility — Stir4Steel". This project has received funding from the Research Fund for Coal and Steel under grant agreement No 101034068.

## References

[1] COMMUNICATION FROM THE COMMISSION TO THE EUROPEAN PARLIAMENT. A Clean Planet for all A European strategic long-term vision for a prosperous, modern, competitive and climate neutral economy, 2018. https://eur-lex.europa.eu/legal-content/EN/TXT/?uri=CELEX:52018DC0773. Lesch, N. Kwiaton, F.B. Klose, Advanced High Strength Steels (AHSS) for Automotive Applications – Tailored Properties by Smart Microstructural Adjustments, Steel Res. Int. 88 (2017) 1700210. https://doi.org/10.1002/srin.201700210

[2] C. Lesch, N. Kwiaton, F.B. Klose, Advanced High Strength Steels (AHSS) for Automotive Applications – Tailored Properties by Smart Microstructural Adjustments, Steel Res. Int. 88 (2017) 1700210. https://doi.org/10.1002/srin.201700210

[3] A.S.M. International, ed., ASM Handbook Volume 1: Properties and Selection: Irons, Steels, and High-Performance Alloys, 10th edition, ASM International, Materials Park, Ohio, 1990.

[4] A.I.H. Committee, ASM Handbook Volume 2: Properties and Selection: Nonferrous Alloys and Special-purpose Materials, ASM International, 1990. https://books.google.de/books?id=wxA7AQAAIAAJ

[5] M.E. Schlesinger, E.M. Mueller, ASM Handbook, Volume 3: Alloy Phase Diagrams, ASM International, Materials Park, Ohio, 1983.

[6] L. Tricarico, R. Spina, D. Sorgente, et al., Effects of heat treatments on mechanical properties of Fe/Al explosion-welded structural transition joints, Mater. Des. 30 (2009) 2693-2700. https://doi.org/10.1016/j.matdes.2008.10.010

[7] R.W. Richards, R.D. Jones, P.D. Clements, et al., Metallurgy of continuous hot dip aluminizing, Int. Mater. Rev. 39 (1994) 191-212. https://doi.org/10.1179/imr.1994.39.5.191.

[8] J. Yang, J.P. Oliveira, Y. Li, et al., Laser techniques for dissimilar joining of aluminum alloys to steels: A critical review, J. Mater. Process. Technol. 301 (2022). https://doi.org/10.1016/j.jmatprotec.2021.117443

[9] J. Chen, B. Shalchi Amirkhiz, R. Zhang, et al., On the Joint Formation and Interfacial Microstructure of Cold Metal Transfer Cycle Step Braze Welding of Aluminum to Steel Butt Joint, Metall. Mater. Trans. A. 51 (2020) 5198-5212. https://doi.org/10.1007/s11661-020-05917-8

[10] G. Chen, J. Liu, X. Shu, et al., Beam scanning effect on properties optimization of thick-plate 2A12 aluminum alloy electron-beam welding joints, Mater. Sci. Eng. A. 744 (2019) 583-592. https://doi.org/10.1016/j.msea.2018.12.034
[11] M.J. Torkamany, S. Tahamtan, J. Sabbaghzadeh, Dissimilar welding of carbon steel to 5754 aluminum alloy by Nd:YAG pulsed laser, Mater. Des. 31 (2010) 458-465.
https://doi.org/10.1016/j.matdes.2009.05.046

[12] G. Qin, Z. Lei, Y. Su, et al., Large spot laser assisted GMA brazing-fusion welding of aluminum alloy to galvanized steel, J. Mater. Process. Technol. 214 (2014) 2684-2692.
https://doi.org/10.1016/j.jmatprotec.2014.06.011

[13] J. Yang, Z. Yu, Y. Li, et al., Influence of alloy elements on microstructure and mechanical properties of Al/steel dissimilar joint by laser welding/brazing, Weld. World. 62 (2018) 427-433.
https://doi.org/10.1007/s40194-017-0540-z

[14] U. Dressler, G. Biallas, und U. Alfaro Mercado, „Friction stir welding of titanium alloy TiAl6V4 to aluminium alloy AA2024-T3“, Mater. Sci. Eng. A, Bd. 526, Nr. 1–2, S. 113–117, Nov. 2009. https://doi.org/10.1016/j.msea.2009.07.006

[15] Y. Kusuda, Honda develops robotized FSW technology to weld steel and aluminum and applied it to a mass-production vehicle, Ind. Robot Int. J. 40 (2013) 208-212.
https://doi.org/10.1108/01439911311309889

[16] L. SELOVA, H. AYDIN, O. TUNCEL, et al., Mechanical Properties of Resistance Spot Welded Three-Sheet Stack Joints of Dissimilar Steels in Different Welding Time, (2019).

[17] T. Matsuda, K. Owada, A. Numata, et al., Influence of interfacial structure on the fracture behavior of friction stir spot welded dissimilar joints, Mater. Sci. Eng. A. 772 (2020) 138743.
https://doi.org/10.1016/j.msea.2019.138743

[18] C. Schilling, J. dos Santos, Method and device for joining at least two adjoining work pieces by friction welding, (2004).

[19] Y. Ding, Z. Shen, A.P. Gerlich, Refill friction stir spot welding of dissimilar aluminum alloy and AlSi coated steel, J. Manuf. Process. 30 (2017) 353-360.
https://doi.org/10.1016/j.jmapro.2017.10.006

[20] DIN EN ISO 14273, Specimen dimensions and procedure for tensile shear testing resistance spot, seam and embossed projection welds. In: DIN, the German Institute for Standardization, 2014.
[11] M.J. Torkamany, S. Tahamtan, J. Sabbaghzadeh, Dissimilar welding of carbon steel to 5754 aluminum alloy by Nd:YAG pulsed laser, Mater. Des. 31 (2010) 458-465.
https://doi.org/10.1016/j.matdes.2009.05.046

[12] G. Qin, Z. Lei, Y. Su, et al., Large spot laser assisted GMA brazing-fusion welding of aluminum alloy to galvanized steel, J. Mater. Process. Technol. 214 (2014) 2684-2692.
https://doi.org/10.1016/j.jmatprotec.2014.06.011

[13] J. Yang, Z. Yu, Y. Li, et al., Influence of alloy elements on microstructure and mechanical properties of Al/steel dissimilar joint by laser welding/brazing, Weld. World. 62 (2018) 427-433.
https://doi.org/10.1007/s40194-017-0540-z

[14] U. Dressler, G. Biallas, und U. Alfaro Mercado, „Friction stir welding of titanium alloy TiAl6V4 to aluminium alloy AA2024-T3“, Mater. Sci. Eng. A, Bd. 526, Nr. 1–2, S. 113–117, Nov. 2009. https://doi.org/10.1016/j.msea.2009.07.006

[15] Y. Kusuda, Honda develops robotized FSW technology to weld steel and aluminum and applied it to a mass-production vehicle, Ind. Robot Int. J. 40 (2013) 208-212.
https://doi.org/10.1108/01439911311309889

[16] L. SELOVA, H. AYDIN, O. TUNCEL, et al., Mechanical Properties of Resistance Spot Welded Three-Sheet Stack Joints of Dissimilar Steels in Different Welding Time, (2019).

[17] T. Matsuda, K. Owada, A. Numata, et al., Influence of interfacial structure on the fracture behavior of friction stir spot welded dissimilar joints, Mater. Sci. Eng. A. 772 (2020) 138743.
https://doi.org/10.1016/j.msea.2019.138743

[18] C. Schilling, J. dos Santos, Method and device for joining at least two adjoining work pieces by friction welding, (2004).

[19] Y. Ding, Z. Shen, A.P. Gerlich, Refill friction stir spot welding of dissimilar aluminum alloy and AlSi coated steel, J. Manuf. Process. 30 (2017) 353-360.
https://doi.org/10.1016/j.jmapro.2017.10.006

[20] DIN EN ISO 14273, Specimen dimensions and procedure for tensile shear testing resistance spot, seam and embossed projection welds. In: DIN, the German Institute for Standardization, 2014.
# Development of new strategies for the mechanical joining of components made of aluminum die casting

KRAUS Christian<fnref n="1" />,ª*, BÖHNKE Max<fnref n="2" />,ᵇ*, DROSSEL Welf-Guntram<fnref n="1" />,ᶜ and
MESCHUT Gerson<fnref n="2" />,ᵈ

**Keywords:** Aluminum Die Casting, Mechanical Joining, Cracking, Heat Treatment, Ductility Testing

**Abstract.** Structures made of aluminum die casting are being used in increasing quantities as well as component sizes for various applications. Due to the size of the components, heat treatment of the cast parts following the casting process is omitted in order to meet dimensional accuracy requirements and reduce production costs. From such an approach, challenges arise with regard to the mechanical joining of these aluminum die-cast structures. On one hand, the absence of heat treatment results in a general decrease in ductility. On the other hand, the increasing size of the components introduces process-related tolerances regarding the quality of the casting, including the presence of air or gas inclusions, and significant variations in ductility within the component. These factors present challenges for mechanical joining technologies, such as the potential risk of crack-related defects during the joining process. For the robust mechanical joining of such materials, the development and validation of suitable joining strategies for aluminum die-cast components is presented in this paper. A preparatory step involving localized heat treatment in the joining area is implemented to enhance the suitability of the casting material for mechanical joining. The objective is to generate an improved ductility state in the aluminum die casting material, enabling crack-free joining through self-pierce riveting. Additionally, the formability of the aluminum die casting material is characterized using a specially developed ductility testing method. This allows the prediction of potential crack-related defects during mechanical joining. The methods described are developed using the AlSi10MnMg material in its as-cast state and applied to the self-pierce riveting process.

**Introduction**

Growing social environmental awareness is forcing the automotive industry to reduce greenhouse gas emissions. To achieve emission-oriented targets, lightweight construction has emerged as a suitable approach. Cast aluminum components are used where high design flexibility is required combined with high structural strength [1]. In recent years, there has been a trend towards increasingly larger structures made of die-cast aluminum, which are replacing complex multi-part structures, such as longitudinal beams, battery housings or rear structures in electric vehicles using giga-casting technologies [2].

Due to the size of the component and the associated time and energy costs, the heat treatment of the cast components that usually follows the casting process is dispensed in order to meet the dimensional accuracy requirements. This in turn results in challenges with regard to the mechanical joining of these die-cast aluminum structures.

<figure><img src="image_1640.png" /></figure>

Content from this work may be used under the terms of the Creative Commons Attribution 3.0 license. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under license by Materials Research Forum LLC.

<footnote n="1">¹Fraunhofer Institute for Machine Tools and Forming Technology IWU, Nöthnitzer Straße 44,
01187 Dresden, Germany</footnote>

<footnote n="2">²Laboratory for material and joining technology LWF, Paderborn University, Pohlweg 47-49,
33098 Paderborn, Germany</footnote>

<footnote>ªchristian.kraus@iwu.fraunhofer.de, ᵇmax.boehnke@lwf.upb.de,</footnote>

<footnote>ᶜwelf-guntram.drossel@iwu.fraunhofer.de, ᵈgerson.meschut@lwf.upb.de</footnote>
# Development of new strategies for the mechanical joining of components made of aluminum die casting

KRAUS Christian<fnref n="1" />,ª*, BÖHNKE Max<fnref n="2" />,ᵇ*, DROSSEL Welf-Guntram<fnref n="1" />,ᶜ and
MESCHUT Gerson<fnref n="2" />,ᵈ

**Keywords:** Aluminum Die Casting, Mechanical Joining, Cracking, Heat Treatment, Ductility Testing

**Abstract.** Structures made of aluminum die casting are being used in increasing quantities as well as component sizes for various applications. Due to the size of the components, heat treatment of the cast parts following the casting process is omitted in order to meet dimensional accuracy requirements and reduce production costs. From such an approach, challenges arise with regard to the mechanical joining of these aluminum die-cast structures. On one hand, the absence of heat treatment results in a general decrease in ductility. On the other hand, the increasing size of the components introduces process-related tolerances regarding the quality of the casting, including the presence of air or gas inclusions, and significant variations in ductility within the component. These factors present challenges for mechanical joining technologies, such as the potential risk of crack-related defects during the joining process. For the robust mechanical joining of such materials, the development and validation of suitable joining strategies for aluminum die-cast components is presented in this paper. A preparatory step involving localized heat treatment in the joining area is implemented to enhance the suitability of the casting material for mechanical joining. The objective is to generate an improved ductility state in the aluminum die casting material, enabling crack-free joining through self-pierce riveting. Additionally, the formability of the aluminum die casting material is characterized using a specially developed ductility testing method. This allows the prediction of potential crack-related defects during mechanical joining. The methods described are developed using the AlSi10MnMg material in its as-cast state and applied to the self-pierce riveting process.

**Introduction**

Growing social environmental awareness is forcing the automotive industry to reduce greenhouse gas emissions. To achieve emission-oriented targets, lightweight construction has emerged as a suitable approach. Cast aluminum components are used where high design flexibility is required combined with high structural strength [1]. In recent years, there has been a trend towards increasingly larger structures made of die-cast aluminum, which are replacing complex multi-part structures, such as longitudinal beams, battery housings or rear structures in electric vehicles using giga-casting technologies [2].

Due to the size of the component and the associated time and energy costs, the heat treatment of the cast components that usually follows the casting process is dispensed in order to meet the dimensional accuracy requirements. This in turn results in challenges with regard to the mechanical joining of these die-cast aluminum structures.

<figure><img src="image_1641.png" /></figure>

Content from this work may be used under the terms of the Creative Commons Attribution 3.0 license. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under license by Materials Research Forum LLC.

<footnote n="1">¹Fraunhofer Institute for Machine Tools and Forming Technology IWU, Nöthnitzer Straße 44,
01187 Dresden, Germany</footnote>

<footnote n="2">²Laboratory for material and joining technology LWF, Paderborn University, Pohlweg 47-49,
33098 Paderborn, Germany</footnote>

<footnote>ªchristian.kraus@iwu.fraunhofer.de, ᵇmax.boehnke@lwf.upb.de,</footnote>

<footnote>ᶜwelf-guntram.drossel@iwu.fraunhofer.de, ᵈgerson.meschut@lwf.upb.de</footnote>
## Self-Pierce riveting of aluminum cast materials

Mechanical joining processes are particularly suitable for connecting the cast aluminum structures to the rest of the body, which consists of different materials due to multi-material designs. Self-pierce riveting has established as a robust and efficient process for joining different types of materials [3]. Due to the brittle material properties of cast aluminum materials, cracks can occur when using the self-pierce riveting technology [4]. Depending on the application, these cracks can reduce the corrosion resistance of the joint [5]. Process and tool parameters can be adjusted to minimize the formation of cracks during the joining process. In [6], it is found that the die contour, especially the die depth, is the main factor in crack initiation during self-pierce riveting in low ductile materials. However, a very influential factor for a specific adjustment of the mechanical properties and thus the reduction of cracking in the cast material is a heat treatment prior to the joining process [7].

Investigations show that the formation of cracks on the joint button can be prevented by suitable heat treatment [7]. The heat treatment primarily affects the energy absorption value and has relatively little influence on the peak force. In addition, the decrease in material strength during heat treatment does not lead to a reduction in joint performance [8].

Due to the immense importance of a prior heat treatment of the aluminum casting material on the one hand and the fluctuating material properties over the entire casting structure on the other hand, two approaches for the robust mechanical joining of such materials are presented in this study. On the one hand, a joint preparation by means of local heat treatment is investigated, which allows a defined ductility state to be generated in the aluminum die casting material and enables crack-free forming joining. Secondly, the forming capability of the aluminum die casting material is to be characterized by using a specially developed ductility testing method. This enables the prediction of possible process-related cracking during mechanical joining.

## Ductility testing methods

Different technological tests are used to determine the ductility of metallic materials during forming or joining processes. However, tensile, compression or deep-drawing tests for material characterization are associated with sample separation from the component and cannot be carried out directly on the component during production. For in-process ductility testing on entire components, it is necessary to use a method that can be carried out locally on the component without significantly damaging the structure.

For the reasons mentioned, a punctual test using a flat punch and a die was developed. This type of test allows characteristic ductility values to be determined directly at relevant joints without damaging the component. The test setup consists of a cylindrical punch and a spring-loaded blank holder on the punch side. On the die side, a tool in the form of a hollow cylinder is attached, see Fig. 1. During the ductility test, the punch presses on the test specimen until a defined drop in force is reached. Conclusions about the ductility of the material to be tested can be drawn from the specific characteristic values such as the maximum punch force $F_{max}$ and the associated punch displacement $s_{Fmax}$ as well as the work applied $W_{Fmax}$ (Fig. 1) [9].
## Self-Pierce riveting of aluminum cast materials

Mechanical joining processes are particularly suitable for connecting the cast aluminum structures to the rest of the body, which consists of different materials due to multi-material designs. Self-pierce riveting has established as a robust and efficient process for joining different types of materials [3]. Due to the brittle material properties of cast aluminum materials, cracks can occur when using the self-pierce riveting technology [4]. Depending on the application, these cracks can reduce the corrosion resistance of the joint [5]. Process and tool parameters can be adjusted to minimize the formation of cracks during the joining process. In [6], it is found that the die contour, especially the die depth, is the main factor in crack initiation during self-pierce riveting in low ductile materials. However, a very influential factor for a specific adjustment of the mechanical properties and thus the reduction of cracking in the cast material is a heat treatment prior to the joining process [7].

Investigations show that the formation of cracks on the joint button can be prevented by suitable heat treatment [7]. The heat treatment primarily affects the energy absorption value and has relatively little influence on the peak force. In addition, the decrease in material strength during heat treatment does not lead to a reduction in joint performance [8].

Due to the immense importance of a prior heat treatment of the aluminum casting material on the one hand and the fluctuating material properties over the entire casting structure on the other hand, two approaches for the robust mechanical joining of such materials are presented in this study. On the one hand, a joint preparation by means of local heat treatment is investigated, which allows a defined ductility state to be generated in the aluminum die casting material and enables crack-free forming joining. Secondly, the forming capability of the aluminum die casting material is to be characterized by using a specially developed ductility testing method. This enables the prediction of possible process-related cracking during mechanical joining.

## Ductility testing methods

Different technological tests are used to determine the ductility of metallic materials during forming or joining processes. However, tensile, compression or deep-drawing tests for material characterization are associated with sample separation from the component and cannot be carried out directly on the component during production. For in-process ductility testing on entire components, it is necessary to use a method that can be carried out locally on the component without significantly damaging the structure.

For the reasons mentioned, a punctual test using a flat punch and a die was developed. This type of test allows characteristic ductility values to be determined directly at relevant joints without damaging the component. The test setup consists of a cylindrical punch and a spring-loaded blank holder on the punch side. On the die side, a tool in the form of a hollow cylinder is attached, see Fig. 1. During the ductility test, the punch presses on the test specimen until a defined drop in force is reached. Conclusions about the ductility of the material to be tested can be drawn from the specific characteristic values such as the maximum punch force $F_{max}$ and the associated punch displacement $s_{Fmax}$ as well as the work applied $W_{Fmax}$ (Fig. 1) [9].
<figure><img src="image_1641.png" /><figcaption>Fig. 1: Experimental setup of the punch indentation test</figcaption></figure>

## Local ductility testing

Tool concept. To characterize the ductility of the untreated and conditioned aluminum die casting material, a punctual punch indentation test is used as described. Specific characteristic values such as the maximum punch force and the associated punch displacement are used to identify correlations regarding the ductility of the casting material. The experimental setup of the test is shown in Fig. 2 a). A stamping tool is used that consists of a punch and a spring-loaded blank holder. Thus, it is possible to vary the punch and die geometry as required.

The tool is implemented in a Zwick Z1486 tensile-compression testing machine, which can be seen in Fig. 2 b). It has an extra-wide testing area that allows complete die-cast components to be tested. It is also characterized by precise displacement control. The punch displacement is measured locally using optical measurement systems and digital image correlation.

<figure><img src="image_1642.png" /><figcaption>Fig. 2: Experimental setup of the punch indentation test</figcaption></figure>

Numerical modeling. In order to identify an optimum tool configuration, the ductility test is first analyzed numerically. For the numerical analysis of the ductility test and to optimize the tool geometries, a 2D model was created in Simufact Forming, which can be seen in Fig. 3. It is a rotationally symmetrical setup with rigid tools. For the validation and parameter study, a material model of a wrought aluminum alloy EN AW-6014 T4 has been implemented in order to avoid the influence factor of varying material properties as it is common for cast materials. The implemented flow curve was determined on the basis of layer compression tests and the extrapolation using the Hockett-Sherby hardening law. A combined friction model consisting of the Coulomb friction model and the friction factor model was used to model the frictional behavior. The Coulomb friction coefficient $\mu$ was set to $\mu = 0.2$ and the friction factor $m = 0.4$. The macro-mechanical