<figure><img src="image_1577.png" /><figcaption>Fig. 1. Load-displacement curved obtained in the Single Lap Shear test for a) co-cured reference aluminium-CFRP specimens without mechanical interlocking joint and b) co-cured aluminium-CFRP specimens with the developed mechanical interlocking joint [8].</figcaption></figure>

However, previous studies on the developed joint were performed at a coupon level in a flat and simple geometry. Yet, in order to scale-up the process, such dissimilar joints need to be able to be formed into complex geometries, if possible, adapting conventional forming processes of the constituent materials to multi-material forming.

Cold forming of aluminium consists of using pressure to force a solid metal blank to deform plastically into the tooling shape. It is performed at temperatures which leave the aluminium microstructure unaffected, does not produce scrap material, results in an acceptable surface finish and it offers better tolerances than hot forming [9]. On the other hand, it requires higher pressure and energy, since aluminium has a poor formability at room temperature [10] and it is known to suffer from abrasive and adhesive wear even at room temperature [11]. The use of lubricants and ceramic PVD (Physical Vapor Deposition) coatings on the tooling delays galling and decreases friction [12,13], being adequate lubrication the most important factor for galling prevention [14].

Feasibility of deep drawing of metal-composite sandwich plates has been assessed in previous literature [15–17]. Conclusions were than when forming the sandwich panels at room temperature, fracture of the aluminium takes place, while an increase of the preheat temperature decreases the severity of failure and increases wrinkling on the aluminium.

This study investigates the viability of cold stamping hybrid aluminium-CFRP prepreg parts with and without the dissimilar joint generated through a punching process, into omega shaped profiles. Additionally, the feasibility of stamping these profiles with and without lubricant application is also considered, since the application of lubricant must be performed carefully in order not to contaminate the uncured epoxy matrix in the CFRP prepreg. Moreover, challenges such as galling, necking, edge fracture and the change in the springback behavior are assessed.

## Materials and methods

Materials. 2,0-mm-thick rolled sheets of AA5754 H111 aluminium alloy were used in this work. This material is commonly used in the automotive industry as a non-heat-treatable aluminium alloy to produce medium strength parts due to its high formability and corrosion resistance.

The selected CFRP was a 0,65-mm-thick twill 2x2 carbon fibre prepreg of 650 g/m² with a 36 % in weight of MTC275 toughened epoxy matrix.

Specimen preparation. Sample processing consisted in three steps: preparation of blanks, joining by punching and eventually stamping of omega components. Different types of specimens were manufactured considering three variables. The resulting test matrix is described in Table 1.

* Position of the CFRP reinforcement: no reinforcement (0), reinforcement inside the omega profile and reinforcement outside the omega profile (o-).
<figure><img src="image_1578.png" /><figcaption>Fig. 1. Load-displacement curved obtained in the Single Lap Shear test for a) co-cured reference aluminium-CFRP specimens without mechanical interlocking joint and b) co-cured aluminium-CFRP specimens with the developed mechanical interlocking joint [8].</figcaption></figure>

However, previous studies on the developed joint were performed at a coupon level in a flat and simple geometry. Yet, in order to scale-up the process, such dissimilar joints need to be able to be formed into complex geometries, if possible, adapting conventional forming processes of the constituent materials to multi-material forming.

Cold forming of aluminium consists of using pressure to force a solid metal blank to deform plastically into the tooling shape. It is performed at temperatures which leave the aluminium microstructure unaffected, does not produce scrap material, results in an acceptable surface finish and it offers better tolerances than hot forming [9]. On the other hand, it requires higher pressure and energy, since aluminium has a poor formability at room temperature [10] and it is known to suffer from abrasive and adhesive wear even at room temperature [11]. The use of lubricants and ceramic PVD (Physical Vapor Deposition) coatings on the tooling delays galling and decreases friction [12,13], being adequate lubrication the most important factor for galling prevention [14].

Feasibility of deep drawing of metal-composite sandwich plates has been assessed in previous literature [15–17]. Conclusions were than when forming the sandwich panels at room temperature, fracture of the aluminium takes place, while an increase of the preheat temperature decreases the severity of failure and increases wrinkling on the aluminium.

This study investigates the viability of cold stamping hybrid aluminium-CFRP prepreg parts with and without the dissimilar joint generated through a punching process, into omega shaped profiles. Additionally, the feasibility of stamping these profiles with and without lubricant application is also considered, since the application of lubricant must be performed carefully in order not to contaminate the uncured epoxy matrix in the CFRP prepreg. Moreover, challenges such as galling, necking, edge fracture and the change in the springback behavior are assessed.

## Materials and methods

Materials. 2,0-mm-thick rolled sheets of AA5754 H111 aluminium alloy were used in this work. This material is commonly used in the automotive industry as a non-heat-treatable aluminium alloy to produce medium strength parts due to its high formability and corrosion resistance.

The selected CFRP was a 0,65-mm-thick twill 2x2 carbon fibre prepreg of 650 g/m² with a 36 % in weight of MTC275 toughened epoxy matrix.

Specimen preparation. Sample processing consisted in three steps: preparation of blanks, joining by punching and eventually stamping of omega components. Different types of specimens were manufactured considering three variables. The resulting test matrix is described in Table 1.

* Position of the CFRP reinforcement: no reinforcement (0), reinforcement inside the omega profile and reinforcement outside the omega profile (o-).
* Lubricant: without (N) and with lubricant (L). In all cases, lubricant was only applied on the aluminium surface of the specimens, meaning in specimens were the CFRP was placed inside the omega profile, the lubricant was placed outside and vice-versa.

* Mechanical joints: specimens with no mechanical joints (0, if it is only aluminium or 1, if it is aluminium and CFRP), specimens with one mechanical joint (2) and specimens with three mechanical joints (3). For the specimens with three mechanical joints, a 20 mm separation space was left between joints.

Four specimens of each type were manufactured. For the same specimen type, the order in which each specimen was stamped was considered, since it is relevant to evaluate adhesive friction and galling of the stamped specimens.

<table><caption>Table 1. Description of the different types of specimens. Inside and outside refer to the side of the omega shaped profile where the CFRP or the lubricant were applied for each specimen type.</caption><thead><tr><th>Specimen</th><th>Aluminium</th><th>CFRP</th><th>Lubricant</th><th>Number of mechanical joints</th></tr></thead><tbody><tr><td>N0</td><td>Yes</td><td>None</td><td>None</td><td>0</td></tr><tr><td>N1</td><td>Yes</td><td>Inside</td><td>None</td><td>0</td></tr><tr><td>N3</td><td>Yes</td><td>Inside</td><td>None</td><td>3</td></tr><tr><td>L0</td><td>Yes</td><td>None</td><td>Outside</td><td>0</td></tr><tr><td>L1</td><td>Yes</td><td>Inside</td><td>Outside</td><td>0</td></tr><tr><td>L2</td><td>Yes</td><td>Inside</td><td>Outside</td><td>1</td></tr><tr><td>L3</td><td>Yes</td><td>Inside</td><td>Outside</td><td>3</td></tr><tr><td>o-L0</td><td>Yes</td><td>None</td><td>Inside</td><td>0</td></tr><tr><td>o-L1</td><td>Yes</td><td>Outside</td><td>Inside</td><td>0</td></tr><tr><td>o-L2</td><td>Yes</td><td>Outside</td><td>Inside</td><td>1</td></tr></tbody></table>

Blank preparation and joining by punching. 125 mm x 250 mm blanks were cut from the aluminium sheet. An uncured CFRP prepreg layer was placed on the blanks, except for specimens N0, L0 and o-L0 (Table 1).

The described mechanical joint was then added to some of the specimens in addition to the adhesive joint coming from the co-bonding of the composite epoxy matrix to the aluminium.

The joint was performed by laying up one uncured CFRP prepreg layer on top of an aluminium sheet and placing a silicone paper layer on top of the CFRP to avoid the uncured epoxy matrix to stick to the punch. Then, the whole system was punched with the CFRP facing the punch side of the set-up (Fig. 2a). The punching was stopped when the aluminum sheet was completely punched but the carbon fibres were not. Instead, the fibres were pressed against the aluminium hole walls, generating a mechanical interlock between both materials (Fig. 2b).

This procedure was performed using a punching tooling mounted on a universal testing machine with a 50 kN loadcell. Punching was performed at a constant speed of 10 mm/min, using a punch with a 10 mm diameter (d<sub>p</sub>, Fig. 2a), as commonly found in the automotive industry [10], and a 0 mm fillet radius (R, Fig. 2a). A die diameter (d<sub>d</sub>, Fig. 2a) of 11,3 mm was used, leading to a clearance of the 24 %, and punching was stopped at a 2,6 mm stroke.
* Lubricant: without (N) and with lubricant (L). In all cases, lubricant was only applied on the aluminium surface of the specimens, meaning in specimens were the CFRP was placed inside the omega profile, the lubricant was placed outside and vice-versa.

* Mechanical joints: specimens with no mechanical joints (0, if it is only aluminium or 1, if it is aluminium and CFRP), specimens with one mechanical joint (2) and specimens with three mechanical joints (3). For the specimens with three mechanical joints, a 20 mm separation space was left between joints.

Four specimens of each type were manufactured. For the same specimen type, the order in which each specimen was stamped was considered, since it is relevant to evaluate adhesive friction and galling of the stamped specimens.

<table><caption>Table 1. Description of the different types of specimens. Inside and outside refer to the side of the omega shaped profile where the CFRP or the lubricant were applied for each specimen type.</caption><thead><tr><th>Specimen</th><th>Aluminium</th><th>CFRP</th><th>Lubricant</th><th>Number of mechanical joints</th></tr></thead><tbody><tr><td>N0</td><td>Yes</td><td>None</td><td>None</td><td>0</td></tr><tr><td>N1</td><td>Yes</td><td>Inside</td><td>None</td><td>0</td></tr><tr><td>N3</td><td>Yes</td><td>Inside</td><td>None</td><td>3</td></tr><tr><td>L0</td><td>Yes</td><td>None</td><td>Outside</td><td>0</td></tr><tr><td>L1</td><td>Yes</td><td>Inside</td><td>Outside</td><td>0</td></tr><tr><td>L2</td><td>Yes</td><td>Inside</td><td>Outside</td><td>1</td></tr><tr><td>L3</td><td>Yes</td><td>Inside</td><td>Outside</td><td>3</td></tr><tr><td>o-L0</td><td>Yes</td><td>None</td><td>Inside</td><td>0</td></tr><tr><td>o-L1</td><td>Yes</td><td>Outside</td><td>Inside</td><td>0</td></tr><tr><td>o-L2</td><td>Yes</td><td>Outside</td><td>Inside</td><td>1</td></tr></tbody></table>

Blank preparation and joining by punching. 125 mm x 250 mm blanks were cut from the aluminium sheet. An uncured CFRP prepreg layer was placed on the blanks, except for specimens N0, L0 and o-L0 (Table 1).

The described mechanical joint was then added to some of the specimens in addition to the adhesive joint coming from the co-bonding of the composite epoxy matrix to the aluminium.

The joint was performed by laying up one uncured CFRP prepreg layer on top of an aluminium sheet and placing a silicone paper layer on top of the CFRP to avoid the uncured epoxy matrix to stick to the punch. Then, the whole system was punched with the CFRP facing the punch side of the set-up (Fig. 2a). The punching was stopped when the aluminum sheet was completely punched but the carbon fibres were not. Instead, the fibres were pressed against the aluminium hole walls, generating a mechanical interlock between both materials (Fig. 2b).

This procedure was performed using a punching tooling mounted on a universal testing machine with a 50 kN loadcell. Punching was performed at a constant speed of 10 mm/min, using a punch with a 10 mm diameter (d<sub>p</sub>, Fig. 2a), as commonly found in the automotive industry [10], and a 0 mm fillet radius (R, Fig. 2a). A die diameter (d<sub>d</sub>, Fig. 2a) of 11,3 mm was used, leading to a clearance of the 24 %, and punching was stopped at a 2,6 mm stroke.
<figure><img src="image_1578.png" /><figcaption>Fig. 2. a) Schematic cross-section of the punching process, b) cured joint from the aluminium sheet side and c) cured joint from the CFRP side.</figcaption></figure>

Stamping. The flat sheets were cold formed into an omega shaped profile (Fig. 3) at 30 Tn using a stamping tooling for hot work steel mounted on a Hydrogarne 150 Tn press. Since joining and stamping take place in an uncured state of the prepreg, delamination and tearing of the CFRP during this process are prevented.

Lubricant was applied prior to stamping. For the lubricated specimens, the RENOFORM 10ALWF 40LB (20 %) lubricant was applied with a roller on the aluminum side of the specimen. However, in specimens featuring mechanical joints, a cautious approach was taken to leave the central 15 mm width unlubricated. This measure aimed to prevent any contamination of the composite epoxy matrix with lubricant. Additionally, a layer of silicon paper was placed atop the CFRP prepreg to mitigate the risk of the epoxy matrix adhering to the stamping tooling.

<figure><img src="image_1579.png" /><figcaption>Fig. 3. Schematic drawing of the stamped omega profile (units in mm).</figcaption></figure>

Joint characterization. After stamping, the joint diameter in the x-direction and in the y-direction (Fig. 4) was measured using a caliper to assess joint deformation.
<figure><img src="image_1579.png" /><figcaption>Fig. 2. a) Schematic cross-section of the punching process, b) cured joint from the aluminium sheet side and c) cured joint from the CFRP side.</figcaption></figure>

Stamping. The flat sheets were cold formed into an omega shaped profile (Fig. 3) at 30 Tn using a stamping tooling for hot work steel mounted on a Hydrogarne 150 Tn press. Since joining and stamping take place in an uncured state of the prepreg, delamination and tearing of the CFRP during this process are prevented.

Lubricant was applied prior to stamping. For the lubricated specimens, the RENOFORM 10ALWF 40LB (20 %) lubricant was applied with a roller on the aluminum side of the specimen. However, in specimens featuring mechanical joints, a cautious approach was taken to leave the central 15 mm width unlubricated. This measure aimed to prevent any contamination of the composite epoxy matrix with lubricant. Additionally, a layer of silicon paper was placed atop the CFRP prepreg to mitigate the risk of the epoxy matrix adhering to the stamping tooling.

<figure><img src="image_1580.png" /><figcaption>Fig. 3. Schematic drawing of the stamped omega profile (units in mm).</figcaption></figure>

Joint characterization. After stamping, the joint diameter in the x-direction and in the y-direction (Fig. 4) was measured using a caliper to assess joint deformation.
<figure><img src="image_1580.png" /><figcaption>Fig. 4. a) L3 specimen after punching and prior to stamping and b) L3 specimen after stamping. Joint deformation is inspected in the stipulated x- and y-directions.</figcaption></figure>

Springback measurement. The flange springback of the specimens was assessed by measuring the θ angle (Fig. 5) with an angle protractor. Thus, higher θ angle correlates to a higher amount of flange springback. A hot stamped Usibor 1500 specimen is used as a no-springback reference (Ref).

<figure><img src="image_1581.png" /><figcaption>Fig. 5. θ Angle measured in order to assess specimens flange springback. R_f corresponds with the radius with the highest friction.</figcaption></figure>

## Results

All specimens could be stamped regardless of the material and joining configuration. However, aluminium and CFRP specimens with three mechanical joints stamped without lubricant (N3) underwent severe necking and edge cracking after the second stamping, being these phenomena more severe at each consecutive stamping (Fig. 6a). These defects were not seen in the analogous specimens stamped with lubricant (L3, Fig. 6b).

When looking at the joint diameter (Fig. 6 and Fig. 7), it can be seen that the diameter in the y-direction remains constant when stamping with (L3, Fig. 6b and Fig. 7d) and without lubrication (N3, Fig. 6a and Fig. 7b) regardless of the number of stamps. However, the joint diameter in the x-direction increases considerably with each consecutive stamping for specimens stamped without lubricant (N3, Fig. 6a and Fig. 7a), meaning the joint is highly deformed in this direction during stamping, and only the first specimen stamped has the same diameter in both x- and y-directions. Nevertheless, the joint diameter in the x-direction of the specimens stamped with lubricant (L3, Fig. 6b and Fig. 7b) remains constant regardless of the number of stamps, and it is the same than the joint diameter in the y-direction.
<figure><img src="image_1581.png" /><figcaption>Fig. 4. a) L3 specimen after punching and prior to stamping and b) L3 specimen after stamping. Joint deformation is inspected in the stipulated x- and y-directions.</figcaption></figure>

Springback measurement. The flange springback of the specimens was assessed by measuring the θ angle (Fig. 5) with an angle protractor. Thus, higher θ angle correlates to a higher amount of flange springback. A hot stamped Usibor 1500 specimen is used as a no-springback reference (Ref).

<figure><img src="image_1582.png" /><figcaption>Fig. 5. θ Angle measured in order to assess specimens flange springback. R_f corresponds with the radius with the highest friction.</figcaption></figure>

## Results

All specimens could be stamped regardless of the material and joining configuration. However, aluminium and CFRP specimens with three mechanical joints stamped without lubricant (N3) underwent severe necking and edge cracking after the second stamping, being these phenomena more severe at each consecutive stamping (Fig. 6a). These defects were not seen in the analogous specimens stamped with lubricant (L3, Fig. 6b).

When looking at the joint diameter (Fig. 6 and Fig. 7), it can be seen that the diameter in the y-direction remains constant when stamping with (L3, Fig. 6b and Fig. 7d) and without lubrication (N3, Fig. 6a and Fig. 7b) regardless of the number of stamps. However, the joint diameter in the x-direction increases considerably with each consecutive stamping for specimens stamped without lubricant (N3, Fig. 6a and Fig. 7a), meaning the joint is highly deformed in this direction during stamping, and only the first specimen stamped has the same diameter in both x- and y-directions. Nevertheless, the joint diameter in the x-direction of the specimens stamped with lubricant (L3, Fig. 6b and Fig. 7b) remains constant regardless of the number of stamps, and it is the same than the joint diameter in the y-direction.
<figure><img src="image_1582.png" /><figcaption>Fig. 6. Stamped aluminium and CFRP specimens with three mechanical joints by stamping order a) stamped without any lubricant and b) stamped with lubricant.</figcaption></figure>

<figure><img src="image_1583.png" /><figcaption>Fig. 7. Joint diameter after stamping in x- and y- direction for specimens with 3 mechanical joints. For specimens N3 (without lubricant) a) diameter in x-direction and b) diameter in y-direction. For specimens L3 (with lubricant) c) diameter in x-direction and d) diameter in y-direction.</figcaption></figure>

Side images of the stamped specimens can be observed in Fig. 8. Visual inspection shows that aluminium without CFRP (N0/Fig. 8a, L0/Fig. 8d and o-L0/Fig. 8h) exhibit a high amount of springback. When adding a CFRP layer in the inside of the omega shaped profile however, the springback is drastically reduced (N1/Fig. 8b, N3/Fig. 8c, L1/Fig. 8e, L2/Fig. 8f and L3/Fig. 8g). Nevertheless, when adding a CFRP layer on the outer part of the profile, such reduction of the springback is not visually observed (o-L1/Fig. 8i and o-L2/Fig. 8j).
<figure><img src="image_1583.png" /><figcaption>Fig. 6. Stamped aluminium and CFRP specimens with three mechanical joints by stamping order a) stamped without any lubricant and b) stamped with lubricant.</figcaption></figure>

<figure><img src="image_1584.png" /><figcaption>Fig. 7. Joint diameter after stamping in x- and y- direction for specimens with 3 mechanical joints. For specimens N3 (without lubricant) a) diameter in x-direction and b) diameter in y-direction. For specimens L3 (with lubricant) c) diameter in x-direction and d) diameter in y-direction.</figcaption></figure>

Side images of the stamped specimens can be observed in Fig. 8. Visual inspection shows that aluminium without CFRP (N0/Fig. 8a, L0/Fig. 8d and o-L0/Fig. 8h) exhibit a high amount of springback. When adding a CFRP layer in the inside of the omega shaped profile however, the springback is drastically reduced (N1/Fig. 8b, N3/Fig. 8c, L1/Fig. 8e, L2/Fig. 8f and L3/Fig. 8g). Nevertheless, when adding a CFRP layer on the outer part of the profile, such reduction of the springback is not visually observed (o-L1/Fig. 8i and o-L2/Fig. 8j).
<figure><img src="image_1584.png" /><figcaption>Fig. 8. Side images of stamped specimens. All images correspond to the 4th stamped specimen of each type. a) N0, b) N1, c) N3, d) L0, e) L1, f) L2, g) L3, h) o-L0, i) o-L1 and j) o-L2.</figcaption></figure>

θ angle measures for each specimen type can be seen in Fig. 9. The measured angles were in agreement with the visual inspection, since aluminium specimens without CFRP (N0/Fig. 9a, L0/Fig. 9b and o-L0/Fig. 9c) and specimens where CFRP had been placed in the outer part of the omega profile (o-L1 and o-L2/Fig. 9c) have a 3° θ angle. However, specimens in where the CFRP was placed inside the omega profile (N1 and N3/Fig. 9a, L1, L2 and L3/Fig. 9b), the θ angle was reduced, implying a decrease of the springback. It can also be seen that in specimens were the CFRP was inside the profile and lubricant was applied, the θ angle increases with the amount of mechanical joints (L1, L2 and L3/Fig. 9b), while this effect was not observed in the analogous specimens where lubricant was not applied (N1 and N3/Fig. 9a).

<figure><img src="image_1585.png" /><figcaption>Fig. 9. θ angle measured for each type of specimens.</figcaption></figure>

Discussion

Results showed that it is possible to cold stamp aluminium with an uncured layer of CFRP prepreg into a complex omega-shaped profile.

However, specimens which were stamped without lubricant (N3) exhibit scratches, necking and edge cracking, after the second stamping operation. The severity of these defects escalated with each successive forming operation, as they were induced by galling. (Fig. 6). Galling is the transfer of aluminium from the stamped part to the tooling surface, followed by hardening of the transferred aluminium by oxidation, work hardening or grain refinement [9]. Thus, in the following stampings, the hardened aluminium increases abruptly the frictional force due to microscopic welding [18], causing scratches in the areas of the part which undergo the most extensive expansion. Therefore, the aluminium hole of the mechanical joint got deformed, becoming wider in the direction of the sheet deformation (x-direction). This caused necking on the perpendicular direction (y-direction)
<figure><img src="image_1585.png" /><figcaption>Fig. 8. Side images of stamped specimens. All images correspond to the 4th stamped specimen of each type. a) N0, b) N1, c) N3, d) L0, e) L1, f) L2, g) L3, h) o-L0, i) o-L1 and j) o-L2.</figcaption></figure>

θ angle measures for each specimen type can be seen in Fig. 9. The measured angles were in agreement with the visual inspection, since aluminium specimens without CFRP (N0/Fig. 9a, L0/Fig. 9b and o-L0/Fig. 9c) and specimens where CFRP had been placed in the outer part of the omega profile (o-L1 and o-L2/Fig. 9c) have a 3° θ angle. However, specimens in where the CFRP was placed inside the omega profile (N1 and N3/Fig. 9a, L1, L2 and L3/Fig. 9b), the θ angle was reduced, implying a decrease of the springback. It can also be seen that in specimens were the CFRP was inside the profile and lubricant was applied, the θ angle increases with the amount of mechanical joints (L1, L2 and L3/Fig. 9b), while this effect was not observed in the analogous specimens where lubricant was not applied (N1 and N3/Fig. 9a).

<figure><img src="image_1586.png" /><figcaption>Fig. 9. θ angle measured for each type of specimens.</figcaption></figure>

Discussion

Results showed that it is possible to cold stamp aluminium with an uncured layer of CFRP prepreg into a complex omega-shaped profile.

However, specimens which were stamped without lubricant (N3) exhibit scratches, necking and edge cracking, after the second stamping operation. The severity of these defects escalated with each successive forming operation, as they were induced by galling. (Fig. 6). Galling is the transfer of aluminium from the stamped part to the tooling surface, followed by hardening of the transferred aluminium by oxidation, work hardening or grain refinement [9]. Thus, in the following stampings, the hardened aluminium increases abruptly the frictional force due to microscopic welding [18], causing scratches in the areas of the part which undergo the most extensive expansion. Therefore, the aluminium hole of the mechanical joint got deformed, becoming wider in the direction of the sheet deformation (x-direction). This caused necking on the perpendicular direction (y-direction)
due to the large amount of strain localized in the narrowest section of the geometry (Fig. 6a), resulting in eventual edge cracking.

The mentioned defects were not present in specimens stamped using lubricant (L3, Fig. 6), since lubrication decreases friction and delays galling [14]. Therefore, even specimens with mechanical joints could be stamped into a complex shape without causing deformation, necking or edge fracture in the vicinity of the aluminium hole, with the use of proper lubrication.

When approaching springback, it was clearly seen that adding an uncured CFRP prepreg layer in the inner part of the omega profile reduced the springback comparing to when stamping aluminium without any CFRP layer. However, when the CFRP prepreg was added on the outer part of the profile, springback appeared to be quite similar to stamping aluminium without CFRP. Clearance between the punch and the die directly affects the drawing force and thus, springback behavior [19]. Nevertheless, the clearance was reduced in the same amount when adding the CFRP, regardless of its position, and is therefore not enough to explain the springback reduction when placing the CFRP in the inner part of the omega. However, this phenomenon can be explained with the fact that the carbon fibres have a high tensile strength and very small deformation (1 %) and the uncured CFRP prepreg exhibits tacky behavior due to the uncured epoxy matrix. Because of that, when placing the CFRP prepreg on the inner part of the omega profile, in the radius with the highest amount of friction (Rf, Fig. 5) the carbon fibres will be in tension and adhered to the aluminium surface due to its tackiness. Therefore, this configuration will restrict the aluminium elastic recovery, thus minimizing flange springback. Nevertheless, when the CFRP prepreg is placed outside the omega profile, the carbon fibres in Rf are compressed. Consequently, elastic recovery of the aluminium in this configuration is not restricted and springback is maintained.

The number of mechanical joints in lubricated specimen also seemed to have an effect on the magnitude of the springback. The higher the number of mechanical joints, the higher the springback. This could be due to the area of CFRP prepreg not adhered to the aluminum in each mechanical joint being more free to move than the rest of the specimen. That involves less restriction to the elastic recovery of the aluminium and, in consequence, larger springback. However, such effect is not observed in the analogous specimens stamped without lubricant, which might be related to the local temperature increase due to friction during stamping, leading to a drop of resin viscosity and tackiness. Clarification of this issue requires further studies.

## Conclusions

The main conclusion of this work is that it is possible to cold stamp aluminium with uncured layers of CFRP prepreg into a complex omega-shaped profile. Nevertheless, specimens featuring mechanical joints must be stamped with lubricant to prevent galling and the subsequent rise in frictional force. This precautionary measure helps avert issues such as necking and edge cracking, particularly after the second forming cycle.

Since two dissimilar materials are being stamped at the same time, material lay-up affects the springback behavior of the stamped part. When the CFRP is placed in the inner part of the omega profile, the springback of the aluminium is highly reduced, while this effect is unseen when the CFRP is found on the outer part of the omega profile.

The number of mechanical joints also influences the springback behavior when lubricant is used during the forming process. The lower the number of mechanical joints, the lower the springback of the stamped part.

## References

[1] J. Fleischer and J. Nieschlag, Introduction to CFRP-metal hybrids for lightweight structures, *Prod. Eng.*, vol. 12, no. 2, pp. 109–111, 2018. https://doi.org/10.1007/s11740-018-0825-0
due to the large amount of strain localized in the narrowest section of the geometry (Fig. 6a), resulting in eventual edge cracking.

The mentioned defects were not present in specimens stamped using lubricant (L3, Fig. 6), since lubrication decreases friction and delays galling [14]. Therefore, even specimens with mechanical joints could be stamped into a complex shape without causing deformation, necking or edge fracture in the vicinity of the aluminium hole, with the use of proper lubrication.

When approaching springback, it was clearly seen that adding an uncured CFRP prepreg layer in the inner part of the omega profile reduced the springback comparing to when stamping aluminium without any CFRP layer. However, when the CFRP prepreg was added on the outer part of the profile, springback appeared to be quite similar to stamping aluminium without CFRP. Clearance between the punch and the die directly affects the drawing force and thus, springback behavior [19]. Nevertheless, the clearance was reduced in the same amount when adding the CFRP, regardless of its position, and is therefore not enough to explain the springback reduction when placing the CFRP in the inner part of the omega. However, this phenomenon can be explained with the fact that the carbon fibres have a high tensile strength and very small deformation (1 %) and the uncured CFRP prepreg exhibits tacky behavior due to the uncured epoxy matrix. Because of that, when placing the CFRP prepreg on the inner part of the omega profile, in the radius with the highest amount of friction (R<sub>f</sub>, Fig. 5) the carbon fibres will be in tension and adhered to the aluminium surface due to its tackiness. Therefore, this configuration will restrict the aluminium elastic recovery, thus minimizing flange springback. Nevertheless, when the CFRP prepreg is placed outside the omega profile, the carbon fibres in R<sub>f</sub> are compressed. Consequently, elastic recovery of the aluminium in this configuration is not restricted and springback is maintained.

The number of mechanical joints in lubricated specimen also seemed to have an effect on the magnitude of the springback. The higher the number of mechanical joints, the higher the springback. This could be due to the area of CFRP prepreg not adhered to the aluminum in each mechanical joint being more free to move than the rest of the specimen. That involves less restriction to the elastic recovery of the aluminium and, in consequence, larger springback. However, such effect is not observed in the analogous specimens stamped without lubricant, which might be related to the local temperature increase due to friction during stamping, leading to a drop of resin viscosity and tackiness. Clarification of this issue requires further studies.

## Conclusions

The main conclusion of this work is that it is possible to cold stamp aluminium with uncured layers of CFRP prepreg into a complex omega-shaped profile. Nevertheless, specimens featuring mechanical joints must be stamped with lubricant to prevent galling and the subsequent rise in frictional force. This precautionary measure helps avert issues such as necking and edge cracking, particularly after the second forming cycle.

Since two dissimilar materials are being stamped at the same time, material lay-up affects the springback behavior of the stamped part. When the CFRP is placed in the inner part of the omega profile, the springback of the aluminium is highly reduced, while this effect is unseen when the CFRP is found on the outer part of the omega profile.

The number of mechanical joints also influences the springback behavior when lubricant is used during the forming process. The lower the number of mechanical joints, the lower the springback of the stamped part.

## References

[1] J. Fleischer and J. Nieschlag, Introduction to CFRP-metal hybrids for lightweight structures, *Prod. Eng.*, vol. 12, no. 2, pp. 109–111, 2018. https://doi.org/10.1007/s11740-018-0825-0
[2] A. Pramanik et al., Joining of carbon fibre reinforced polymer (CFRP) composites and aluminium alloys – A review, *Compos. Part A Appl. Sci. Manuf.*, vol. 101, pp. 1–29, 2017. https://doi.org/10.1016/j.compositesa.2017.06.007

[3] Y. Ouyang and C. Chen, Research advances in the mechanical joining process for fiber reinforced plastic composites, *Compos. Struct.*, vol. 296, no. March, p. 115906, 2022. https://doi.org/10.1016/j.compstruct.2022.115906

[4] A. Galińska, Mechanical joining of fibre reinforced polymer composites to metals—a review. Part I: Bolted joining, *Polymers (Basel)*, vol. 12, no. 10, pp. 1–48, 2020. https://doi.org/10.3390/polym12102252

[5] G. Di Franco, L. Fratini, and A. Pasta, Analysis of the mechanical performance of hybrid (SPR/bonded) single-lap joints between CFRP panels and aluminum blanks, *Int. J. Adhes. Adhes.*, vol. 41, pp. 24–32, 2013. https://doi.org/10.1016/j.ijadhadh.2012.10.008

[6] F. Lambiase and D. C. Ko, Feasibility of mechanical clinching for joining aluminum AA6082-T6 and Carbon Fiber Reinforced Polymer sheets, *Mater. Des.*, vol. 107, pp. 341–352, 2016. https://doi.org/10.1016/j.matdes.2016.06.061

[7] A. Galińska and C. Galiński, Mechanical joining of fibre reinforced polymer composites to metals-A review. Part II: Riveting, clinching, non-adhesive form-locked joints, pin and loop joining, *Polymers (Basel)*, vol. 12, no. 8, 2020. https://doi.org/10.3390/POLYM12081681

[8] N. Latorre, D. Casellas, and J. Costa, A punching process to join metal sheets and fibre reinforced polymer composites by mechanical interlocking, *Compos. Part A*, vol. 175, no. July, p. 107744, 2023. https://doi.org/10.1016/j.compositesa.2023.107744

[9] J. Heinrichs and S. Jacobson, The influence from shape and size of tool surface defects on the occurrence of galling in cold forming of aluminium, *Wear*, vol. 271, no. 9–10, pp. 2517–2524, 2011. https://doi.org/10.1016/j.wear.2011.01.077

[10] K. Lange, Handbook of Metal Forming, *J. Appl. Metalwork.*, vol. 4, no. 2, p. 188, 1986. https://doi.org/10.1007/BF02834383

[11] J. Pujante, L. Pelcastre, M. Vilaseca, D. Casellas, and B. Prakash, Investigations into wear and galling mechanism of aluminium alloy-tool steel tribopair at different temperatures, *Wear*, vol. 308, no. 1–2, pp. 193–198, 2013. https://doi.org/10.1016/j.wear.2013.06.015

[12] M. Vilaseca, S. Molas, and D. Casellas, High temperature tribological behaviour of tool steels during sliding against aluminium, *Wear*, vol. 272, no. 1, pp. 105–109, 2011. https://doi.org/10.1016/j.wear.2011.07.007

[13] M. Vilaseca, J. Pujante, G. Ramírez, and D. Casellas, Investigation into adhesive wear of PVD coated and uncoated hot stamping production tools, *Wear*, vol. 308, no. 1–2, pp. 148–154, 2013. https://doi.org/10.1016/j.wear.2013.07.003

[14] J. Heinrichs and S. Jacobson, Laboratory test simulation of aluminium cold forming - influence from PVD tool coatings on the tendency to galling, *Surf. Coatings Technol.*, vol. 204, no. 21–22, pp. 3606–3613, 2010. https://doi.org/10.1016/j.surfcoat.2010.04.025

[15] S. F. Hwang and Y. R. Li, Deep Drawing Behavior of Metal-Composite Sandwich Plates, *Materials (Basel)*, vol. 15, no. 19, 2022. https://doi.org/10.3390/ma15196612

[16] J. Gresham, W. Cantwell, M. J. Cardew-Hall, P. Compston, and S. Kalyanasundaram, Drawing behaviour of metal-composite sandwich structures, *Compos. Struct.*, vol. 75, no. 1–4, pp. 305–312, 2006. https://doi.org/10.1016/j.compstruct.2006.04.010

[17] T. Heggemann and W. Homberg, Deep drawing of fiber metal laminates for automotive lightweight structures, *Compos. Struct.*, vol. 216, no. December 2018, pp. 53–57, 2019.
[2] A. Pramanik et al., Joining of carbon fibre reinforced polymer (CFRP) composites and aluminium alloys – A review, *Compos. Part A Appl. Sci. Manuf.*, vol. 101, pp. 1–29, 2017. https://doi.org/10.1016/j.compositesa.2017.06.007

[3] Y. Ouyang and C. Chen, Research advances in the mechanical joining process for fiber reinforced plastic composites, *Compos. Struct.*, vol. 296, no. March, p. 115906, 2022. https://doi.org/10.1016/j.compstruct.2022.115906

[4] A. Galińska, Mechanical joining of fibre reinforced polymer composites to metals—a review. Part I: Bolted joining, *Polymers (Basel)*, vol. 12, no. 10, pp. 1–48, 2020. https://doi.org/10.3390/polym12102252

[5] G. Di Franco, L. Fratini, and A. Pasta, Analysis of the mechanical performance of hybrid (SPR/bonded) single-lap joints between CFRP panels and aluminum blanks, *Int. J. Adhes. Adhes.*, vol. 41, pp. 24–32, 2013. https://doi.org/10.1016/j.ijadhadh.2012.10.008

[6] F. Lambiase and D. C. Ko, Feasibility of mechanical clinching for joining aluminum AA6082-T6 and Carbon Fiber Reinforced Polymer sheets, *Mater. Des.*, vol. 107, pp. 341–352, 2016. https://doi.org/10.1016/j.matdes.2016.06.061

[7] A. Galińska and C. Galiński, Mechanical joining of fibre reinforced polymer composites to metals-A review. Part II: Riveting, clinching, non-adhesive form-locked joints, pin and loop joining, *Polymers (Basel)*, vol. 12, no. 8, 2020. https://doi.org/10.3390/POLYM12081681

[8] N. Latorre, D. Casellas, and J. Costa, A punching process to join metal sheets and fibre reinforced polymer composites by mechanical interlocking, *Compos. Part A*, vol. 175, no. July, p. 107744, 2023. https://doi.org/10.1016/j.compositesa.2023.107744

[9] J. Heinrichs and S. Jacobson, The influence from shape and size of tool surface defects on the occurrence of galling in cold forming of aluminium, *Wear*, vol. 271, no. 9–10, pp. 2517–2524, 2011. https://doi.org/10.1016/j.wear.2011.01.077

[10] K. Lange, Handbook of Metal Forming, *J. Appl. Metalwork.*, vol. 4, no. 2, p. 188, 1986. https://doi.org/10.1007/BF02834383

[11] J. Pujante, L. Pelcastre, M. Vilaseca, D. Casellas, and B. Prakash, Investigations into wear and galling mechanism of aluminium alloy-tool steel tribopair at different temperatures, *Wear*, vol. 308, no. 1–2, pp. 193–198, 2013. https://doi.org/10.1016/j.wear.2013.06.015

[12] M. Vilaseca, S. Molas, and D. Casellas, High temperature tribological behaviour of tool steels during sliding against aluminium, *Wear*, vol. 272, no. 1, pp. 105–109, 2011. https://doi.org/10.1016/j.wear.2011.07.007

[13] M. Vilaseca, J. Pujante, G. Ramírez, and D. Casellas, Investigation into adhesive wear of PVD coated and uncoated hot stamping production tools, *Wear*, vol. 308, no. 1–2, pp. 148–154, 2013. https://doi.org/10.1016/j.wear.2013.07.003

[14] J. Heinrichs and S. Jacobson, Laboratory test simulation of aluminium cold forming - influence from PVD tool coatings on the tendency to galling, *Surf. Coatings Technol.*, vol. 204, no. 21–22, pp. 3606–3613, 2010. https://doi.org/10.1016/j.surfcoat.2010.04.025

[15] S. F. Hwang and Y. R. Li, Deep Drawing Behavior of Metal-Composite Sandwich Plates, *Materials (Basel)*, vol. 15, no. 19, 2022. https://doi.org/10.3390/ma15196612

[16] J. Gresham, W. Cantwell, M. J. Cardew-Hall, P. Compston, and S. Kalyanasundaram, Drawing behaviour of metal-composite sandwich structures, *Compos. Struct.*, vol. 75, no. 1–4, pp. 305–312, 2006. https://doi.org/10.1016/j.compstruct.2006.04.010

[17] T. Heggemann and W. Homberg, Deep drawing of fiber metal laminates for automotive lightweight structures, *Compos. Struct.*, vol. 216, no. December 2018, pp. 53–57, 2019.
https://doi.org/10.1016/j.compstruct.2019.02.047

[18] K. Dohda, M. Yamamoto, C. Hu, L. Dubar, and K. F. Ehmann, Galling phenomena in metal forming, *Friction*, vol. 9, no. 4, pp. 665-685, 2021. https://doi.org/10.1007/s40544-020-0430-z

[19] A. Wang *et al.*, Springback analysis of AA5754 after hot stamping: experiments and FE modelling, *Int. J. Adv. Manuf. Technol.*, vol. 89, no. 5-8, pp. 1339-1352, 2017. https://doi.org/10.1007/s00170-016-9166-3
https://doi.org/10.1016/j.compstruct.2019.02.047

[18] K. Dohda, M. Yamamoto, C. Hu, L. Dubar, and K. F. Ehmann, Galling phenomena in metal forming, *Friction*, vol. 9, no. 4, pp. 665-685, 2021. https://doi.org/10.1007/s40544-020-0430-z

[19] A. Wang *et al.*, Springback analysis of AA5754 after hot stamping: experiments and FE modelling, *Int. J. Adv. Manuf. Technol.*, vol. 89, no. 5-8, pp. 1339-1352, 2017. https://doi.org/10.1007/s00170-016-9166-3
# Non-rotationally symmetric joints
– Mechanisms and load bearing capacity

DEVULAPALLY Deekshith Reddy¹,ᵃ*, MARTIN Sven¹,ᵇ and TRÖSTER Thomas¹,ᶜ

¹Paderborn University, Chair of Automotive Lightweight Design, Warburger Str. 100, 33098 Paderborn, Germany

ªdeekshith.reddy.devulapally@uni-paderborn.de, ᵇsven.martin@uni-paderborn.de,

ᶜthomas.troester@uni-paderborn.de

**Keywords:** Non-Rotationally Symmetric Joints, Load Distribution, Joint Design, Anisotropic Joints

**Abstract.** Traditionally, joints are cylindrical and rotationally symmetric. In the present study, non-rotationally symmetric joints are used for joining steel and Glass mat-reinforced thermoplastic sheets (GMT). In addition, the study also analyzes the impact of non-rotational symmetric joint rotation on the load-bearing capacity. Single lap joint specimens were fabricated using the In-Mold assembly technique for joining steel sheets with GMT. Tensile shear tests were performed on different orientations of the joint geometry, and it was observed that changing the joint orientation influences the load-bearing capacity. The joints are constitutively modeled using beam elements and the influence of joint rotation on load distribution is examined through a static simulation study.

**Introduction**

The joining of manufactured components presents a significant hurdle in any industry, especially in the case of automobile body construction, where thousands of joints are made using different joining processes. Reducing or optimizing the joints according to loading conditions can reduce the number of joints, saving many resources[1]. Due to their ease of joining, most joints are typically rotationally symmetric; nevertheless, using these traditional joints might reduce the degree of design freedom. Non-rotationally symmetric joints can have an additional degree of freedom by changing the orientation of the joint geometry in a specific direction. Since they are rotationally non-symmetric, they have good torsional resistance and instead of using two rotationally symmetric joints one non-rotationally symmetric joint can be used. This helps in reducing the number of joints which also helps in saving production time.

The knowledge of load distribution and the type of loads acting on joints is adequately understood because of the complexity of loading. In most cases, the joints are either designed with a higher factor of safety or their number is increased, which increases the cost, weight, and time. The study conducted by [2] on joint loading in global bending and torsion in an automotive body in white (BIW) shows that most of the joints are shear-loaded. This is because more than 70% of the joints are lap joints. The following study conducted by [3] on loading the BIW with different real-life maneuvers also shows that joints are mainly shear-loaded and suggests a strong potential for joints with non-rotationally symmetric geometry.

In addition, substantial work on the load distribution among joints with rotationally symmetric geometry exists in the literature [4]. According to this study, when arranged in a line, the load distribution among the joints is uneven; the joints near the edges take most of the load, comparatively less load. The investigation by [5] on load transfer in multi-spot ultrasonic welded joints in thermoplastic composites also confirms the uneven load distribution among the joints in a line. Furthermore, the study conducted by [6] on load transfer in multi-fastener joints of metal

<figure><img src="image_1586.png" /></figure>
# Non-rotationally symmetric joints
– Mechanisms and load bearing capacity

DEVULAPALLY Deekshith Reddy¹,ᵃ*, MARTIN Sven¹,ᵇ and TRÖSTER Thomas¹,ᶜ

¹Paderborn University, Chair of Automotive Lightweight Design, Warburger Str. 100, 33098 Paderborn, Germany

ᵃdeekshith.reddy.devulapally@uni-paderborn.de, ᵇsven.martin@uni-paderborn.de,

ᶜthomas.troester@uni-paderborn.de

**Keywords:** Non-Rotationally Symmetric Joints, Load Distribution, Joint Design, Anisotropic Joints

**Abstract.** Traditionally, joints are cylindrical and rotationally symmetric. In the present study, non-rotationally symmetric joints are used for joining steel and Glass mat-reinforced thermoplastic sheets (GMT). In addition, the study also analyzes the impact of non-rotational symmetric joint rotation on the load-bearing capacity. Single lap joint specimens were fabricated using the In-Mold assembly technique for joining steel sheets with GMT. Tensile shear tests were performed on different orientations of the joint geometry, and it was observed that changing the joint orientation influences the load-bearing capacity. The joints are constitutively modeled using beam elements and the influence of joint rotation on load distribution is examined through a static simulation study.

**Introduction**

The joining of manufactured components presents a significant hurdle in any industry, especially in the case of automobile body construction, where thousands of joints are made using different joining processes. Reducing or optimizing the joints according to loading conditions can reduce the number of joints, saving many resources[1]. Due to their ease of joining, most joints are typically rotationally symmetric; nevertheless, using these traditional joints might reduce the degree of design freedom. Non-rotationally symmetric joints can have an additional degree of freedom by changing the orientation of the joint geometry in a specific direction. Since they are rotationally non-symmetric, they have good torsional resistance and instead of using two rotationally symmetric joints one non-rotationally symmetric joint can be used. This helps in reducing the number of joints which also helps in saving production time.

The knowledge of load distribution and the type of loads acting on joints is adequately understood because of the complexity of loading. In most cases, the joints are either designed with a higher factor of safety or their number is increased, which increases the cost, weight, and time. The study conducted by [2] on joint loading in global bending and torsion in an automotive body in white (BIW) shows that most of the joints are shear-loaded. This is because more than 70% of the joints are lap joints. The following study conducted by [3] on loading the BIW with different real-life maneuvers also shows that joints are mainly shear-loaded and suggests a strong potential for joints with non-rotationally symmetric geometry.

In addition, substantial work on the load distribution among joints with rotationally symmetric geometry exists in the literature [4]. According to this study, when arranged in a line, the load distribution among the joints is uneven; the joints near the edges take most of the load, comparatively less load. The investigation by [5] on load transfer in multi-spot ultrasonic welded joints in thermoplastic composites also confirms the uneven load distribution among the joints in a line. Furthermore, the study conducted by [6] on load transfer in multi-fastener joints of metal

<figure><img src="image_1587.png" /></figure>
and composite laminates suggests that joints with increased stiffness can carry more loads than others and help in even load distribution in a line. It was also suggested that increasing the diameter of the bolts in the middle of the configuration can help in the even distribution of loads because of the increase in stiffness with increasing the bolt geometry. The present study investigates the effect of non-rotationally symmetric joints and changing the joints' orientation on even load distribution. The current investigation focused on non-rotationally symmetric joints produced from a combination of steel and GMT, specifically designed in an elliptical shape. The joints were arranged in three orientations by changing their orientation angles. The effect of orientation on joint stiffness and load-bearing capacity was investigated, and the impact of the locking mechanism of the joint was also in focus. Joints with and without interlocking and failure mechanisms of joint combinations were examined by considering tensile shear test properties as well. At the end of the study, preliminary numerical studies were carried out with joints in an array, and the load distribution was analyzed. Moreover, it was also considered whether changes in the orientation of joints affect the load distribution among other joints.

## Method

Tensile shear specimens are produced by joining steel sheets with fiber composite material using a non-rotationally symmetric form-fit joint. The joining process consists of a two-step operation. In the first step, the two joining members to be joined are prepared. Hole flanging is employed to prepare the steel sheet, followed by cutting the GMT into the necessary dimensions and subsequently heating it in the oven. In the second step, both the joining members are stacked upon each other in a heated mold and pressed together. Different variations of the joint are produced by rotating the joint geometry. The joints were subjected to shear testing via a universal testing machine, and the load vs displacement data were analyzed to observe the bearing capacity of the joints.

### Materials.

The thermoplastic composite material used in this study is a chopped fiberglass mat reinforced thermoplastic (GMT) with randomly oriented glass fibers in a mineral-filled polypropylene (PP) matrix. It is commercially known as S153A248-M1 and is procured from Mitsubishi Chemical Advanced Materials Composites AG with a glass fiber content of 52%. The sheet metal used is HCT590X (1.0941), a dual-phase steel with remarkably high tensile strength and good cold-forming capacity. The material is zinc-coated and procured from Salzgitter Flachstahl GmbH. The properties of both materials are tabulated in Table 1. Laser cutting is used for preparing sheet metal parts, offering high precision and good cutting surface quality. The dimensions of the laser-cut specimens can be seen in Fig. 1.

### Hole-Flanging.

Hole-flanging, or collar drawing, is a tension-compression metal forming technique described in the German norm DIN 8584-1 punch-and-die setup with a blank holder is used for forming flanges around holes in sheet metal components. The blank cut hole is clamped by a blank holder and is pushed into the drawing die by a rigid punch, which forms a flange in the shape of the die around the hole. This process is used extensively in the automotive industry for lightweight applications [8]. In the present study, elliptical hole flanges are formed in the sheet metal component and are used to examine the joining of metal with GMT. The flange acts as an anchor and can bear load in the joint. Flanges were produced in three orientations (0°, 45°, and 90°) and were investigated in the current study. The process is schematically represented in Fig. 2 below.
and composite laminates suggests that joints with increased stiffness can carry more loads than others and help in even load distribution in a line. It was also suggested that increasing the diameter of the bolts in the middle of the configuration can help in the even distribution of loads because of the increase in stiffness with increasing the bolt geometry. The present study investigates the effect of non-rotationally symmetric joints and changing the joints' orientation on even load distribution. The current investigation focused on non-rotationally symmetric joints produced from a combination of steel and GMT, specifically designed in an elliptical shape. The joints were arranged in three orientations by changing their orientation angles. The effect of orientation on joint stiffness and load-bearing capacity was investigated, and the impact of the locking mechanism of the joint was also in focus. Joints with and without interlocking and failure mechanisms of joint combinations were examined by considering tensile shear test properties as well. At the end of the study, preliminary numerical studies were carried out with joints in an array, and the load distribution was analyzed. Moreover, it was also considered whether changes in the orientation of joints affect the load distribution among other joints.

## Method

Tensile shear specimens are produced by joining steel sheets with fiber composite material using a non-rotationally symmetric form-fit joint. The joining process consists of a two-step operation. In the first step, the two joining members to be joined are prepared. Hole flanging is employed to prepare the steel sheet, followed by cutting the GMT into the necessary dimensions and subsequently heating it in the oven. In the second step, both the joining members are stacked upon each other in a heated mold and pressed together. Different variations of the joint are produced by rotating the joint geometry. The joints were subjected to shear testing via a universal testing machine, and the load vs displacement data were analyzed to observe the bearing capacity of the joints.

### Materials.

The thermoplastic composite material used in this study is a chopped fiberglass mat reinforced thermoplastic (GMT) with randomly oriented glass fibers in a mineral-filled polypropylene (PP) matrix. It is commercially known as S153A248-M1 and is procured from Mitsubishi Chemical Advanced Materials Composites AG with a glass fiber content of 52%. The sheet metal used is HCT590X (1.0941), a dual-phase steel with remarkably high tensile strength and good cold-forming capacity. The material is zinc-coated and procured from Salzgitter Flachstahl GmbH. The properties of both materials are tabulated in Table 1. Laser cutting is used for preparing sheet metal parts, offering high precision and good cutting surface quality. The dimensions of the laser-cut specimens can be seen in Fig. 1.

### Hole-Flanging.

Hole-flanging, or collar drawing, is a tension-compression metal forming technique described in the German norm DIN 8584-1 punch-and-die setup with a blank holder is used for forming flanges around holes in sheet metal components. The blank cut hole is clamped by a blank holder and is pushed into the drawing die by a rigid punch, which forms a flange in the shape of the die around the hole. This process is used extensively in the automotive industry for lightweight applications [8]. In the present study, elliptical hole flanges are formed in the sheet metal component and are used to examine the joining of metal with GMT. The flange acts as an anchor and can bear load in the joint. Flanges were produced in three orientations (0°, 45°, and 90°) and were investigated in the current study. The process is schematically represented in Fig. 2 below.
<figure><img src="image_1587.png" /><figcaption>Figure. 1 Laser-cut sheet metal with dimensions in three orientations.</figcaption></figure>

<table><caption>Table 1. Material details</caption><thead><tr><td>Material</td><td>Thickness [mm]</td><td>Tensile strength [MPa]</td><td>Young's modulus [GPa]</td><td>Density [g/cm³]</td></tr></thead><tbody><tr><td>HCT590X- Steel</td><td>1.5</td><td>590</td><td>210</td><td>7.85</td></tr><tr><td>S153A248-M1- Composite</td><td>4.8</td><td>145</td><td>86</td><td>1.40</td></tr></tbody></table>

<figure><img src="image_1588.png" /><figcaption>Figure. 2 Process of sheet metal hole-flanging process.</figcaption></figure>

In mold assembly.

The two materials were joined by In-Mold Assembly (IMA) of hybrid components using compressive molding [9]. GMT laminate is cut into the required dimensions by a plate shear and is heated in the convection oven for 10 mins at 250°C. Steel specimen with a hole flange is inserted into the molding tool, which is maintained at 100°C; the hot GMT specimens are then immediately transferred and stacked onto the metal specimen, which is pressed via a mechanical press (Synchropress 1M-300). Higher compressive forces, combined with good flowable properties of the GMT material, ensure that the material is uniformly distributed and ideally takes the shape of the mold. The process flow of the joining method is shown schematically in Fig. 3. Two laminates are stacked one upon the other to increase the thickness and strength of the composite part. During the compressive molding process, the heated thermoplastic composite material flows into the flange, forming a secured joint. The fiber flow and the interlock can be observed by a cut-section
<figure><img src="image_1588.png" /><figcaption>Figure. 1 Laser-cut sheet metal with dimensions in three orientations.</figcaption></figure>

<table><caption>Table 1. Material details</caption><thead><tr><td>Material</td><td>Thickness [mm]</td><td>Tensile strength [MPa]</td><td>Young's modulus [GPa]</td><td>Density [g/cm³]</td></tr></thead><tbody><tr><td>HCT590X- Steel</td><td>1.5</td><td>590</td><td>210</td><td>7.85</td></tr><tr><td>S153A248-M1- Composite</td><td>4.8</td><td>145</td><td>86</td><td>1.40</td></tr></tbody></table>

<figure><img src="image_1589.png" /><figcaption>Figure. 2 Process of sheet metal hole-flanging process.</figcaption></figure>

In mold assembly.

The two materials were joined by In-Mold Assembly (IMA) of hybrid components using compressive molding [9]. GMT laminate is cut into the required dimensions by a plate shear and is heated in the convection oven for 10 mins at 250°C. Steel specimen with a hole flange is inserted into the molding tool, which is maintained at 100°C; the hot GMT specimens are then immediately transferred and stacked onto the metal specimen, which is pressed via a mechanical press (Synchropress 1M-300). Higher compressive forces, combined with good flowable properties of the GMT material, ensure that the material is uniformly distributed and ideally takes the shape of the mold. The process flow of the joining method is shown schematically in Fig. 3. Two laminates are stacked one upon the other to increase the thickness and strength of the composite part. During the compressive molding process, the heated thermoplastic composite material flows into the flange, forming a secured joint. The fiber flow and the interlock can be observed by a cut-section
view of the specimen in Fig. 5. The dimensions of the lap shear specimen and the process flow are described in Fig. 4. In the experimental investigations conducted, two types of flanges were utilized: one without edge interlocking and the other with interlocking. The influence of interlocking, joint fracture, and load-bearing capability in various orientations are investigated and compared.

<figure><img src="image_1589.png" /><figcaption>Figure. 3 Process of joining through in mold assembly through compressive molding.</figcaption></figure>

<figure><img src="image_1590.png" /><figcaption>Figure. 4 Joined specimen with dimensions.</figcaption></figure>

<figure><img src="image_1591.png" /><figcaption>Figure. 5 Cut section of specimens A) Specimen without edge interlock. B) Specimen with edge interlock.</figcaption></figure>

Finite Element Simulation

The finite element simulation comprises two steps. In the first step, the joint stiffness derived from experimental data is utilized to calibrate individual joints. Subsequently, in the second phase, calibrated single joints with various orientation angles are employed to simulate multiple joints arranged in a line. For the simulation, an implicit solver of the Software Ansys LSDYNA and
view of the specimen in Fig. 5. The dimensions of the lap shear specimen and the process flow are described in Fig. 4. In the experimental investigations conducted, two types of flanges were utilized: one without edge interlocking and the other with interlocking. The influence of interlocking, joint fracture, and load-bearing capability in various orientations are investigated and compared.

<figure><img src="image_1590.png" /><figcaption>Figure. 3 Process of joining through in mold assembly through compressive molding.</figcaption></figure>

<figure><img src="image_1591.png" /><figcaption>Figure. 4 Joined specimen with dimensions.</figcaption></figure>

<figure><img src="image_1592.png" /><figcaption>Figure. 5 Cut section of specimens A) Specimen without edge interlock. B) Specimen with edge interlock.</figcaption></figure>

Finite Element Simulation

The finite element simulation comprises two steps. In the first step, the joint stiffness derived from experimental data is utilized to calibrate individual joints. Subsequently, in the second phase, calibrated single joints with various orientation angles are employed to simulate multiple joints arranged in a line. For the simulation, an implicit solver of the Software Ansys LSDYNA and
Hyper mesh preprocessing software is used for meshing. The model is created by considering the dimensions of the joint overlap area in the experiments. In the calibration phase, only the overlap of two joining members is considered and is simulated. The two joining members are represented by fully integrated shell elements (elform=16), and thickness values are given as the thickness of steel and GMT. The joint is modeled using discrete beams (beam formulation=6) with a beam length equal to the total thickness of the joint. Discrete beam elements are 1-D elements with six degrees of freedom generally used for representing and calibrating joints and fasteners. This beam element has a local coordinate system (r, s, t); the two beam nodes connecting the two shell members are aligned along the r-direction. The third node of the beam can be aligned in s or t-direction according to the direction of stiffness (see Fig. 7). Beam elements are connected to the shell members using the constrained nodal rigid body (CNRB) constraint. Fig. 6 shows the FE model with a single joint. The steel member is subjected to full constraint, while the composite member experiences a prescribed displacement of 0.4 mm in the x-direction over a duration of 8.04 seconds, as indicated by experimental observations.

<figure><img src="image_1592.png" /><figcaption>Figure. 6 FE Model setup for single joint calibration</figcaption></figure>

The stiffness values are calculated regarding the experimental results by linear curve fitting in Origin pro software. The stiffness values are defined as an input in the material data card of the discrete beam element *MAT_LINEAR_ELASTIC DISCRETE BEAM. The material data of the joining members are also obtained from the material data sheets and implemented in the simulation using the *MAT_ELASTIC card. More information about the mentioned keywords can be found in the LS-DYNA keyword user's manual [10].

Contact definition between the joining members is represented using the contact card *AUTOMATIC Surface TO Surface MORTAR. The composite member is considered master, and the steel member is determined as an enslaved person with a coefficient of static friction value of 0.1. Since no pull-out tension tests are experimentally conducted, a normal stiffness value of 10000 N/mm is defined as a constant value in the calibration of joints. The average stiffness values for three different joint rotations can be seen in the following Table 3. Since no rotation takes place during testing, the rotational degree of freedom is restricted in all three directions. All the joints were calibrated for a displacement of 0.4 mm and a time of 8.04 seconds, which was taken from the experimental results, and the same conditions are also used for simulating multijoint setup.

<table><caption>Table 3 Stiffness values for different joint orientation</caption><thead><tr><td>Rotation Angle</td><td>Average Stiffness [N/mm]</td><td>Standard Deviation</td></tr></thead><tbody><tr><td>0°</td><td>11550</td><td>142.94</td></tr><tr><td>45°</td><td>11300</td><td>294.90</td></tr><tr><td>90°</td><td>12000</td><td>893.81</td></tr></tbody></table>
Hyper mesh preprocessing software is used for meshing. The model is created by considering the dimensions of the joint overlap area in the experiments. In the calibration phase, only the overlap of two joining members is considered and is simulated. The two joining members are represented by fully integrated shell elements (elform=16), and thickness values are given as the thickness of steel and GMT. The joint is modeled using discrete beams (beam formulation=6) with a beam length equal to the total thickness of the joint. Discrete beam elements are 1-D elements with six degrees of freedom generally used for representing and calibrating joints and fasteners. This beam element has a local coordinate system (r, s, t); the two beam nodes connecting the two shell members are aligned along the r-direction. The third node of the beam can be aligned in s or t-direction according to the direction of stiffness (see Fig. 7). Beam elements are connected to the shell members using the constrained nodal rigid body (CNRB) constraint. Fig. 6 shows the FE model with a single joint. The steel member is subjected to full constraint, while the composite member experiences a prescribed displacement of 0.4 mm in the x-direction over a duration of 8.04 seconds, as indicated by experimental observations.

<figure><img src="image_1593.png" /><figcaption>Figure. 6 FE Model setup for single joint calibration</figcaption></figure>

The stiffness values are calculated regarding the experimental results by linear curve fitting in Origin pro software. The stiffness values are defined as an input in the material data card of the discrete beam element *MAT_LINEAR_ELASTIC DISCRETE BEAM. The material data of the joining members are also obtained from the material data sheets and implemented in the simulation using the *MAT_ELASTIC card. More information about the mentioned keywords can be found in the LS-DYNA keyword user's manual [10].

Contact definition between the joining members is represented using the contact card *AUTOMATIC Surface TO Surface MORTAR. The composite member is considered master, and the steel member is determined as an enslaved person with a coefficient of static friction value of 0.1. Since no pull-out tension tests are experimentally conducted, a normal stiffness value of 10000 N/mm is defined as a constant value in the calibration of joints. The average stiffness values for three different joint rotations can be seen in the following Table 3. Since no rotation takes place during testing, the rotational degree of freedom is restricted in all three directions. All the joints were calibrated for a displacement of 0.4 mm and a time of 8.04 seconds, which was taken from the experimental results, and the same conditions are also used for simulating multijoint setup.

<table><caption>Table 3 Stiffness values for different joint orientation</caption><thead><tr><td>Rotation Angle</td><td>Average Stiffness [N/mm]</td><td>Standard Deviation</td></tr></thead><tbody><tr><td>0°</td><td>11550</td><td>142.94</td></tr><tr><td>45°</td><td>11300</td><td>294.90</td></tr><tr><td>90°</td><td>12000</td><td>893.81</td></tr></tbody></table>
<figure><img src="image_1593.png" /><figcaption>Figure.7 FE model with multi-joint setup</figcaption></figure>

## Results and discussion

The non-rotationally symmetric joint specimens were tested for lap shear strength using the universal testing machine MTS Criterion C45 with a testing speed of 3 mm/s. For each orientation of the joint, three tests were conducted to reduce the errors and abnormal results. The steel member of the specimen is fixed, and the composite part is stretched. All the specimens were tested until failure, and the load vs displacement data was acquired using the integrated load cell in the testing machine. Fig. 8 shows the stages of the specimen during the testing. Fig. 9 compares the load vs displacement data for three orientations. The linear part of the graph represents the elastic region. After this stage, plastic deformation occurs, and the joint tends to rotate because of the secondary bending. The two joining members separate as the specimen is further stretched, reducing the joint's load-bearing capacity. It can be observed that a change in the joint's rotation significantly affects the maximum load-bearing capacity of the joint. Joints with 45° orientation bear the most negligible load, and joints with 0° orientation show maximum load-bearing capacity. The peak load-bearing capacity of the joints is given in Table 4 below. One more critical observation from the graph is the displacement value at which the maximum load is seen. Joints with 0° orientation show maximum load at a displacement value, which is more than the other two orientations. From this, it can be concluded that a change in orientation affects not only the load-bearing capacity but also the displacement at which the joint is bearing the maximum load. The following Fig.10 presents a comparison of joints with and without edge interlocking. It can be noticed that the joints with interlock perform noticeably lower compared to those without interlocking. The reason for this might be because of the length of the metal flange in the composite member. Since only a small portion of the metal flange is in the composite member, joints with interlock were fractured early, thus reducing the maximum load-bearing capacity.

<table><caption>Table 4 Maximum load with change in joint orientation</caption><thead><tr><td>Orientation angle</td><td>Maximum load [N] (Without interlock)</td><td>Maximum load [N] (With interlock)</td></tr></thead><tbody><tr><td>0°</td><td>7959.56</td><td>7033.04</td></tr><tr><td>45°</td><td>5972.35</td><td>5843.82</td></tr><tr><td>90°</td><td>7216.25</td><td>5833.39</td></tr></tbody></table>
<figure><img src="image_1594.png" /><figcaption>Figure.7 FE model with multi-joint setup</figcaption></figure>

## Results and discussion

The non-rotationally symmetric joint specimens were tested for lap shear strength using the universal testing machine MTS Criterion C45 with a testing speed of 3 mm/s. For each orientation of the joint, three tests were conducted to reduce the errors and abnormal results. The steel member of the specimen is fixed, and the composite part is stretched. All the specimens were tested until failure, and the load vs displacement data was acquired using the integrated load cell in the testing machine. Fig. 8 shows the stages of the specimen during the testing. Fig. 9 compares the load vs displacement data for three orientations. The linear part of the graph represents the elastic region. After this stage, plastic deformation occurs, and the joint tends to rotate because of the secondary bending. The two joining members separate as the specimen is further stretched, reducing the joint's load-bearing capacity. It can be observed that a change in the joint's rotation significantly affects the maximum load-bearing capacity of the joint. Joints with 45° orientation bear the most negligible load, and joints with 0° orientation show maximum load-bearing capacity. The peak load-bearing capacity of the joints is given in Table 4 below. One more critical observation from the graph is the displacement value at which the maximum load is seen. Joints with 0° orientation show maximum load at a displacement value, which is more than the other two orientations. From this, it can be concluded that a change in orientation affects not only the load-bearing capacity but also the displacement at which the joint is bearing the maximum load. The following Fig.10 presents a comparison of joints with and without edge interlocking. It can be noticed that the joints with interlock perform noticeably lower compared to those without interlocking. The reason for this might be because of the length of the metal flange in the composite member. Since only a small portion of the metal flange is in the composite member, joints with interlock were fractured early, thus reducing the maximum load-bearing capacity.

<table><caption>Table 4 Maximum load with change in joint orientation</caption><thead><tr><td>Orientation angle</td><td>Maximum load [N] (Without interlock)</td><td>Maximum load [N] (With interlock)</td></tr></thead><tbody><tr><td>0°</td><td>7959.56</td><td>7033.04</td></tr><tr><td>45°</td><td>5972.35</td><td>5843.82</td></tr><tr><td>90°</td><td>7216.25</td><td>5833.39</td></tr></tbody></table>
<figure><img src="image_1594.png" /><figcaption>Figure. 8 Secondary bending of the metal plate during shear testing</figcaption></figure>

The load distribution study is conducted using the calibrated values obtained from the single-joint setup and is subsequently applied to a configuration comprising four joints arranged in a line. First, all the joints with similar rotation are considered and are simulated in the load distribution study. Given that these joints are calibrated primarily based on stiffness, the load distribution study does not reveal any noticeable distinction in even load distribution, as there are minimal fluctuations in stiffness values. The reason for uneven load distribution is still not known but it might be the cause of using different thicknesses of joining members. Fig. 11 shows that the initial joint bears most of the load, while the subsequent joints carry comparatively lesser loads. In addition, several joint orientation combinations were studied to evaluate possible differences in the load distribution. Since there is no large difference in stiffness between different orientations a large change in load distribution has not been observed. Making use of this finding, the method of modifying stiffness values to provide the best possible load distribution across joints is investigated. It was found that a decrease in the first two joints' stiffness combined with an increase in the last two joints' stiffness had advantageous results for load distribution (see Fig. 11). Joints with similar orientation angles are only examined for stiffness optimization study, the additional freedom of changing the joint orientation adds one more degree of freedom in this case. This shows that the usage of joints with different stiffness can help distribute the load evenly in multi-joint setups. Further investigations into the impact of alterations in the size and geometry of non-rotationally symmetric joints on stiffness would advance our understanding. Moreover, examining the influence of inter-joint distances would provide additional insights.
<figure><img src="image_1595.png" /><figcaption>Figure. 8 Secondary bending of the metal plate during shear testing</figcaption></figure>

The load distribution study is conducted using the calibrated values obtained from the single-joint setup and is subsequently applied to a configuration comprising four joints arranged in a line. First, all the joints with similar rotation are considered and are simulated in the load distribution study. Given that these joints are calibrated primarily based on stiffness, the load distribution study does not reveal any noticeable distinction in even load distribution, as there are minimal fluctuations in stiffness values. The reason for uneven load distribution is still not known but it might be the cause of using different thicknesses of joining members. Fig. 11 shows that the initial joint bears most of the load, while the subsequent joints carry comparatively lesser loads. In addition, several joint orientation combinations were studied to evaluate possible differences in the load distribution. Since there is no large difference in stiffness between different orientations a large change in load distribution has not been observed. Making use of this finding, the method of modifying stiffness values to provide the best possible load distribution across joints is investigated. It was found that a decrease in the first two joints' stiffness combined with an increase in the last two joints' stiffness had advantageous results for load distribution (see Fig. 11). Joints with similar orientation angles are only examined for stiffness optimization study, the additional freedom of changing the joint orientation adds one more degree of freedom in this case. This shows that the usage of joints with different stiffness can help distribute the load evenly in multi-joint setups. Further investigations into the impact of alterations in the size and geometry of non-rotationally symmetric joints on stiffness would advance our understanding. Moreover, examining the influence of inter-joint distances would provide additional insights.
<figure><img src="image_1595.png" /><figcaption>Figure. 9 Load vs. displacement graph for three orientations without the interlock.</figcaption></figure>

<figure><img src="image_1596.png" /><figcaption>Figure. 10 Comparison of load vs displacement graph for three orientations</figcaption></figure>
<figure><img src="image_1596.png" /><figcaption>Figure. 9 Load vs. displacement graph for three orientations without the interlock.</figcaption></figure>

<figure><img src="image_1597.png" /><figcaption>Figure. 10 Comparison of load vs displacement graph for three orientations</figcaption></figure>
<figure><img src="image_1597.png" /><figcaption>Figure. 11 Load distribution in multi-joint setup using non-rotationally symmetric joints. All 0 corresponds to all joints oriented in 0 orientation and All 0_opt corresponds to joints with optimized stiffness.</figcaption></figure>

**Conclusion**

Non-rotationally symmetric joints were used to join steel and composite materials utilizing the In-Mold assembly technique. The joints were made with and without interlock configuration and tested for load-bearing efficiency—joints with interlocking mechanisms performed remarkably lower than those without one. Changing the orientation angle of the joints affects the maximum load-bearing characteristics. It was observed that joints with 0° orientations show higher load-bearing capacity in comparison to other orientations. The load distribution study by arranging multiple joints in a line shows the need for stiffer joints in even load distribution. Further experimental investigations on multi-joint setups using non-rotationally symmetric joints and changing the joints' shape and inter-joint distance can help understand these joints better.

**Acknowledgment**

This research was funded by the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation), project number 418701707 - TRR 285 phase -2 (Sub-project B01). The authors gratefully acknowledge the financial support. The authors express their gratitude to the Paderborn Center for Parallel Computing (PC2) for generously providing the computational resources essential for conducting this research.

**References**

[1] A. B. Ryberg and L. Nilsson, "Spot weld reduction methods for automotive structures," *Structural and Multidisciplinary Optimization*, vol. 53, no. 4, pp. 923-934, Apr. 2016. https://doi.org/10.1007/S00158-015-1355-4/TABLES/4

[2] S. Martin, J. Schütte, C. Bäumler, W. Sextro, and T. Tröster, "Identification of joints for a load-adapted shape in a body in white using steady state vehicle simulations," *Forces in Mechanics*, vol. 6, p. 100065, Feb. 2022. https://doi.org/10.1016/J.FINMEC.2021.100065

[3] S. Martin, A. A. Camberg, and T. Tröster, "Probability Distribution of Joint Point Loadings in Car Body Structures under Global Bending and Torsion," *Procedia Manuf*, vol. 47, pp. 419-424, Jan. 2020. https://doi.org/10.1016/J.PROMFG.2020.04.324

[4] Q. Zhao et al., "Insights into the structural design strategies of multi-spot ultrasonic welded joints in thermoplastic composites: A finite element analysis," *Compos Struct*, vol. 299, p. 115996, Nov. 2022. https://doi.org/10.1016/J.COMPSTRUCT.2022.115996
<figure><img src="image_1598.png" /><figcaption>Figure. 11 Load distribution in multi-joint setup using non-rotationally symmetric joints. All 0 corresponds to all joints oriented in 0 orientation and All 0_opt corresponds to joints with optimized stiffness.</figcaption></figure>

**Conclusion**

Non-rotationally symmetric joints were used to join steel and composite materials utilizing the In-Mold assembly technique. The joints were made with and without interlock configuration and tested for load-bearing efficiency—joints with interlocking mechanisms performed remarkably lower than those without one. Changing the orientation angle of the joints affects the maximum load-bearing characteristics. It was observed that joints with 0° orientations show higher load-bearing capacity in comparison to other orientations. The load distribution study by arranging multiple joints in a line shows the need for stiffer joints in even load distribution. Further experimental investigations on multi-joint setups using non-rotationally symmetric joints and changing the joints' shape and inter-joint distance can help understand these joints better.

**Acknowledgment**

This research was funded by the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation), project number 418701707 - TRR 285 phase -2 (Sub-project B01). The authors gratefully acknowledge the financial support. The authors express their gratitude to the Paderborn Center for Parallel Computing (PC2) for generously providing the computational resources essential for conducting this research.

**References**

[1] A. B. Ryberg and L. Nilsson, "Spot weld reduction methods for automotive structures," *Structural and Multidisciplinary Optimization*, vol. 53, no. 4, pp. 923-934, Apr. 2016. https://doi.org/10.1007/S00158-015-1355-4/TABLES/4

[2] S. Martin, J. Schütte, C. Bäumler, W. Sextro, and T. Tröster, "Identification of joints for a load-adapted shape in a body in white using steady state vehicle simulations," *Forces in Mechanics*, vol. 6, p. 100065, Feb. 2022. https://doi.org/10.1016/J.FINMEC.2021.100065

[3] S. Martin, A. A. Camberg, and T. Tröster, "Probability Distribution of Joint Point Loadings in Car Body Structures under Global Bending and Torsion," *Procedia Manuf*, vol. 47, pp. 419-424, Jan. 2020. https://doi.org/10.1016/J.PROMFG.2020.04.324

[4] Q. Zhao et al., "Insights into the structural design strategies of multi-spot ultrasonic welded joints in thermoplastic composites: A finite element analysis," *Compos Struct*, vol. 299, p. 115996, Nov. 2022. https://doi.org/10.1016/J.COMPSTRUCT.2022.115996
[5] J. Ekh and J. Schön, “Load transfer in multirow, single shear, composite-to-aluminium lap joints,” *Compos Sci Technol*, vol. 66, no. 7–8, pp. 875–885, Jun. 2006. https://doi.org/10.1016/J.COMPSCITECH.2005.08.015

[6] J. Ekh and J. Schön, “Finite element modeling and optimization of load transfer in multi-fastener joints using structural elements,” *Compos Struct*, vol. 82, no. 2, pp. 245–256, Jan. 2008. https://doi.org/10.1016/J.COMPSTRUCT.2007.01.005

[7] “DIN 8584-5 - Manufacturing processes forming under combination of tensile and compressive conditions - Part 5: Forming by raising; Classification, subdivision, terms and definitions | GlobalSpec.” Accessed: Dec. 14, 2023. [Online]. Available: https://standards.globalspec.com/std/430962/DIN%208584-5

[8] Y. Dewang, R. Purohit, and N. Tenguria, “A study on sheet metal hole-flanging process,” *Mater Today Proc*, vol. 4, no. 4, pp. 5421–5428, Jan. 2017. https://doi.org/10.1016/J.MATPR.2017.05.053

[9] T. Stallmeister and T. Tröster, “In-Mold-Assembly of Hybrid Bending Structures by Compression Molding,” *Key Eng Mater*, vol. 926 KEM, pp. 1457–1467, 2022. https://doi.org/10.4028/P-5FXP53

[10] “LS-DYNA ® KEYWORD USER’S MANUAL VOLUME II Material Models LIVERMORE SOFTWARE TECHNOLOGY CORPORATION (LSTC),” 2012, Accessed: Jan. 12, 2024. [Online]. Available: www.lstc.com
[5] J. Ekh and J. Schön, “Load transfer in multirow, single shear, composite-to-aluminium lap joints,” *Compos Sci Technol*, vol. 66, no. 7–8, pp. 875–885, Jun. 2006. https://doi.org/10.1016/J.COMPSCITECH.2005.08.015

[6] J. Ekh and J. Schön, “Finite element modeling and optimization of load transfer in multi-fastener joints using structural elements,” *Compos Struct*, vol. 82, no. 2, pp. 245–256, Jan. 2008. https://doi.org/10.1016/J.COMPSTRUCT.2007.01.005

[7] “DIN 8584-5 - Manufacturing processes forming under combination of tensile and compressive conditions - Part 5: Forming by raising; Classification, subdivision, terms and definitions | GlobalSpec.” Accessed: Dec. 14, 2023. [Online]. Available: https://standards.globalspec.com/std/430962/DIN%208584-5

[8] Y. Dewang, R. Purohit, and N. Tenguria, “A study on sheet metal hole-flanging process,” *Mater Today Proc*, vol. 4, no. 4, pp. 5421–5428, Jan. 2017. https://doi.org/10.1016/J.MATPR.2017.05.053

[9] T. Stallmeister and T. Tröster, “In-Mold-Assembly of Hybrid Bending Structures by Compression Molding,” *Key Eng Mater*, vol. 926 KEM, pp. 1457–1467, 2022. https://doi.org/10.4028/P-5FXP53

[10] “LS-DYNA ® KEYWORD USER’S MANUAL VOLUME II Material Models LIVERMORE SOFTWARE TECHNOLOGY CORPORATION (LSTC),” 2012, Accessed: Jan. 12, 2024. [Online]. Available: www.lstc.com
# Assessments of staked hybrid joints made by studs 3D printed at different manufacturing conditions

CONTE Romina<fnref n="1" />, a*, GAGLIARDI Francesco<fnref n="1" />, b, AMBROGIO Giuseppina<fnref n="1" />, c and
FILICE Luigino<fnref n="1" />, d

**Keywords:** Processing Technologies, Staking Process, Hybrid Structures, Polymer-Metal Joining, Mechanical Testing

**Abstract.** The integration of polymers and metals in engineering applications has led to the development of hybrid joints that combine the distinct advantages of both materials. The staking process is a manufacturing technique, which involves the joining or the securing components through the controlled deformation of materials. It involves forming a permanent and strong connection by reshaping or displacing material, often using force or pressure without the need of additional fasteners, like screws or adhesives. This study presents a comprehensive experimental analysis of a specific hybrid joint performed by staking and by using a polymeric 3D printed partner and a steel plate. The influence of the printer machine and the related printing speed on the joint strength is discussed and the findings demonstrate a strong correlation between this variable and the mechanical properties of the hybrid joints.

## Introduction

Designers and engineers are involved in the selection of lighter, recyclable, and sustainable materials due to increasingly stringent environmental policies. In fact, the lightweight design, and specifically in the transportation sector, responds to the need to reduce weight, aiming to limit emissions of harmful substances to the environment and climate.

The integration of polymers and metals in engineering applications has led to the development of hybrid joints that combine the distinct advantages of both materials and play a fundamental role in lightweighting, particularly in automotive and aerospace, as well as in the medical field for the development of implants and prosthetics, and in civil engineering for bridge construction. They promote sustainable production while simultaneously allowing the creation of robust structures.

To create hybrid joints, it is necessary to optimize both material selection and geometry simultaneously. Hybrid joints are created by combining different materials with diverse properties to achieve the desired performance. Therefore, despite the advantages of using lightweight materials, the union of metal alloys and composites poses a significant challenge due to their different chemical and physical properties. On one hand, metals possess key properties such as high strength, high toughness, high thermal conductivity, and good stability at high temperatures. On the other hand, plastics are characterized by lightness, high corrosion resistance, and excellent formability. However, due to the different properties, problems may arise during the joining process, as well as in terms of the structural integrity of the component [1].

Recent studies have concentrated on the assembly of multimaterial structures using joining technologies distinct from traditional methods, including mechanical, bonding, and adhesive techniques. The analysis encompassed examination of failure modes and mechanical characteristics of the resulting hybrid joints.

Abibe et al. [2] experimented the Injection Clinching Joining (ICJ) as an alternative joining method for connecting polyamide 6,6 reinforced with 30% short glass fibres with aluminium alloy,

<figure><img src="image_1598.png" /></figure>

<footnote n="1">¹University of Calabria, Ponte P. Bucci - Cubo 45C, 87036 Arcavacata di Rende (CS) - Italy</footnote>

<footnote>aromina.conte@unical.it, bfrancesco.gagliardi@unical.it, cgiuseppina.ambrogio@unical.it,</footnote>

<footnote>dluigino.filice@unical.it</footnote>
# Assessments of staked hybrid joints made by studs 3D printed at different manufacturing conditions

CONTE Romina<fnref n="1" />,ᵃ*, GAGLIARDI Francesco<fnref n="1" />,ᵇ, AMBROGIO Giuseppina<fnref n="1" />,ᶜ and
FILICE Luigino<fnref n="1" />,ᵈ

**Keywords:** Processing Technologies, Staking Process, Hybrid Structures, Polymer-Metal Joining, Mechanical Testing

**Abstract.** The integration of polymers and metals in engineering applications has led to the development of hybrid joints that combine the distinct advantages of both materials. The staking process is a manufacturing technique, which involves the joining or the securing components through the controlled deformation of materials. It involves forming a permanent and strong connection by reshaping or displacing material, often using force or pressure without the need of additional fasteners, like screws or adhesives. This study presents a comprehensive experimental analysis of a specific hybrid joint performed by staking and by using a polymeric 3D printed partner and a steel plate. The influence of the printer machine and the related printing speed on the joint strength is discussed and the findings demonstrate a strong correlation between this variable and the mechanical properties of the hybrid joints.

## Introduction

Designers and engineers are involved in the selection of lighter, recyclable, and sustainable materials due to increasingly stringent environmental policies. In fact, the lightweight design, and specifically in the transportation sector, responds to the need to reduce weight, aiming to limit emissions of harmful substances to the environment and climate.

The integration of polymers and metals in engineering applications has led to the development of hybrid joints that combine the distinct advantages of both materials and play a fundamental role in lightweighting, particularly in automotive and aerospace, as well as in the medical field for the development of implants and prosthetics, and in civil engineering for bridge construction. They promote sustainable production while simultaneously allowing the creation of robust structures.

To create hybrid joints, it is necessary to optimize both material selection and geometry simultaneously. Hybrid joints are created by combining different materials with diverse properties to achieve the desired performance. Therefore, despite the advantages of using lightweight materials, the union of metal alloys and composites poses a significant challenge due to their different chemical and physical properties. On one hand, metals possess key properties such as high strength, high toughness, high thermal conductivity, and good stability at high temperatures. On the other hand, plastics are characterized by lightness, high corrosion resistance, and excellent formability. However, due to the different properties, problems may arise during the joining process, as well as in terms of the structural integrity of the component [1].

Recent studies have concentrated on the assembly of multimaterial structures using joining technologies distinct from traditional methods, including mechanical, bonding, and adhesive techniques. The analysis encompassed examination of failure modes and mechanical characteristics of the resulting hybrid joints.

Abibe et al. [2] experimented the Injection Clinching Joining (ICJ) as an alternative joining method for connecting polyamide 6,6 reinforced with 30% short glass fibres with aluminium alloy,

<figure><img src="image_1599.png" /></figure>

<footnote n="1">¹University of Calabria, Ponte P. Bucci - Cubo 45C, 87036 Arcavacata di Rende (CS) - Italy</footnote>

<footnote>ªromina.conte@unical.it, ᵇfrancesco.gagliardi@unical.it, ᶜgiuseppina.ambrogio@unical.it,</footnote>

<footnote>ᵈluigino.filice@unical.it</footnote>
2024-T351, used for automotive components and engineering structures, respectively. The microstructures and the failure modes of overlap joints were assessed and related to the process conditions, and the mechanical performance was considered satisfactory.

Liu et al. [3] designed a new technique, named friction lap welding (FLW) for direct joining AZ31B Mg alloy polyamide 6, aiming at analysing the effect of the welding parameters on the bubble formation and at improving the strength of the hybrid joints. Recently the authors demonstrated a robust bond at interface, by using aluminum alloy and polyamide 66 (PA66) surpassing mere mechanical interlocking, as evidenced by XPS measurements [4].

Abibe et al. [5] introduced for the first time, the innovative Friction-based Injection Clinching Joining technique (F-ICJ). The authors formed joints between polyetherimide (PEI) and aluminium alloy 6082-T6, provided details on the microstructure and alterations in local properties, and demonstrated that the process represent a promising joining technology for polymer-metal hybrid structures, also for some advantages related to the absence of gas o fume and to the general less environmental impact.

Huang et al. [6] investigated a novel technique named Friction Filling Staking Joining (FFSJ) for joining 6082-T6 aluminium alloy and poly propylene polymer sheets. The mechanical performance obtained, really close to the state-of-the-art welding, demonstrate the potential of this technique.

Liu et al. [7] presented a high-speed friction lap welding (FLW) technique for directly joining of aluminium to polymer. The influence of the welding temperature and speed was analysed thought X-ray observation and lap-shear tests. The results revealed that the bond strength was consistent irrespective of the cited investigated parameters.

Huang et al. [8] provided a review on the laser joining technology of polymers and metals, highlighting the impact of some factors, such as the bubbles morphology or the depth of molten pool, on the joining strengths.

Considering the growing interest in directly connecting metal to polymer material, this paper focused on the staking process, another interesting joining technique, which find application in several fields (e.g. aerospace and automotive) and allows the creation of a mechanical interlocking through the application of pressure and deformation, and without the use of additional fastening elements, such as screws [9]. Various forms of staking exist (e.g. cold staking, heated-tool staking, hot-air staking, ultrasonic staking and infrared/laser staking), and the selection of a particular method depends mainly on the materials to be processed, the distinct application, and the specific joint requirement.

A general scheme of the staking process has been described by Härtel et al. [10]. It consists, first, in the preparation and positioning of the components to be joined (Fig. 1 – a) and, then thought a staking tool, such as a punch, pressure in applied on the rivet pin (Fig. 1 – b), therefore this latter is deformed, becoming a rivet-like element, and a mechanical interlocking is created (Fig. 1 – c).

<figure><img src="image_1599.png" /><figcaption>Fig. 1 - Staking process principle [10]</figcaption></figure>
2024-T351, used for automotive components and engineering structures, respectively. The microstructures and the failure modes of overlap joints were assessed and related to the process conditions, and the mechanical performance was considered satisfactory.

Liu et al. [3] designed a new technique, named friction lap welding (FLW) for direct joining AZ31B Mg alloy polyamide 6, aiming at analysing the effect of the welding parameters on the bubble formation and at improving the strength of the hybrid joints. Recently the authors demonstrated a robust bond at interface, by using aluminum alloy and polyamide 66 (PA66) surpassing mere mechanical interlocking, as evidenced by XPS measurements [4].

Abibe et al. [5] introduced for the first time, the innovative Friction-based Injection Clinching Joining technique (F-ICJ). The authors formed joints between polyetherimide (PEI) and aluminium alloy 6082-T6, provided details on the microstructure and alterations in local properties, and demonstrated that the process represent a promising joining technology for polymer-metal hybrid structures, also for some advantages related to the absence of gas o fume and to the general less environmental impact.

Huang et al. [6] investigated a novel technique named Friction Filling Staking Joining (FFSJ) for joining 6082-T6 aluminium alloy and poly propylene polymer sheets. The mechanical performance obtained, really close to the state-of-the-art welding, demonstrate the potential of this technique.

Liu et al. [7] presented a high-speed friction lap welding (FLW) technique for directly joining of aluminium to polymer. The influence of the welding temperature and speed was analysed thought X-ray observation and lap-shear tests. The results revealed that the bond strength was consistent irrespective of the cited investigated parameters.

Huang et al. [8] provided a review on the laser joining technology of polymers and metals, highlighting the impact of some factors, such as the bubbles morphology or the depth of molten pool, on the joining strengths.

Considering the growing interest in directly connecting metal to polymer material, this paper focused on the staking process, another interesting joining technique, which find application in several fields (e.g. aerospace and automotive) and allows the creation of a mechanical interlocking through the application of pressure and deformation, and without the use of additional fastening elements, such as screws [9]. Various forms of staking exist (e.g. cold staking, heated-tool staking, hot-air staking, ultrasonic staking and infrared/laser staking), and the selection of a particular method depends mainly on the materials to be processed, the distinct application, and the specific joint requirement.

A general scheme of the staking process has been described by Härtel et al. [10]. It consists, first, in the preparation and positioning of the components to be joined (Fig. 1 – a) and, then thought a staking tool, such as a punch, pressure in applied on the rivet pin (Fig. 1 – b), therefore this latter is deformed, becoming a rivet-like element, and a mechanical interlocking is created (Fig. 1 – c).

<figure><img src="image_1600.png" /><figcaption>Fig. 1 - Staking process principle [10]</figcaption></figure>
This study assesses the Hot-Air-Cold-Staking (HACS) variant, aiming at testing the joining feasibility between polymer and metal, evaluating the impact of the key process parameters on the rivet pin geometry and on the resulting joint strength assessed by lap-shear test.

## Materials and method

The materials used to create the joint are the polylactic acid (PLA) and steel C40.

PLA is a biomaterial widely used for industrial applications, replacing conventional petroleum-based polymers, therefore it is considered environmentally friendly. Furthermore, it is a biodegradable aliphatic polyester that can be produced from renewable raw materials. It is a thermoplastic polymer, with high tensile strength and rigidity [11]. Additionally, it has a better thermal workability compared to other biopolymers and can be processed using additive manufacturing techniques. In fact, it is one of the most commonly used materials in the Fused Deposition Modeling (FDM) process [12] due to its ease of printing, low smell emissions and biodegradability. Steel C40, on the other hand, is classified as a non-alloy steel and is a high-quality carbon structural steel applied for various industrial applications owing to its strength, ductility, toughness and good machinability.

The physical and mechanical properties of both materials are displayed in Table 1.

<table><caption>Table 1 – Physical and mechanical properties of the investigated materials</caption><thead><tr><th></th><th>PLA</th><th>Steel C40</th></tr></thead><tbody><tr><td>Density (ρ) [g/cm³]</td><td>1,21-1,25</td><td>7,85</td></tr><tr><td>Tensile strength (σ) [MPa]</td><td>21-60</td><td>650-800</td></tr><tr><td>Young's modulus (E) [GPa]</td><td>0,35-3,5</td><td>210</td></tr><tr><td>Elongation (ε) [%]</td><td>2,5-6</td><td>20-25</td></tr><tr><td>Glass transition temperature (Tg) [°C]</td><td>45-60</td><td>-</td></tr><tr><td>Melting temperature (Tm) [°C]</td><td>150-162</td><td>-</td></tr></tbody></table>

Fig. 2 - displays the geometries of the metal and polymeric partners, based on the ASTM: D5961 standard, which provides details on the execution of the lap shear test (Fig. 2 - b). The steel C40 sheet has been drilled according to the suggested hole size, while the PLA sheet has been produced by additive manufacturing and it was characterized by a rivet pin height of 8 mm.
This study assesses the Hot-Air-Cold-Staking (HACS) variant, aiming at testing the joining feasibility between polymer and metal, evaluating the impact of the key process parameters on the rivet pin geometry and on the resulting joint strength assessed by lap-shear test.

## Materials and method

The materials used to create the joint are the polylactic acid (PLA) and steel C40.

PLA is a biomaterial widely used for industrial applications, replacing conventional petroleum-based polymers, therefore it is considered environmentally friendly. Furthermore, it is a biodegradable aliphatic polyester that can be produced from renewable raw materials. It is a thermoplastic polymer, with high tensile strength and rigidity [11]. Additionally, it has a better thermal workability compared to other biopolymers and can be processed using additive manufacturing techniques. In fact, it is one of the most commonly used materials in the Fused Deposition Modeling (FDM) process [12] due to its ease of printing, low smell emissions and biodegradability. Steel C40, on the other hand, is classified as a non-alloy steel and is a high-quality carbon structural steel applied for various industrial applications owing to its strength, ductility, toughness and good machinability.

The physical and mechanical properties of both materials are displayed in Table 1.

<table><caption>Table 1 – Physical and mechanical properties of the investigated materials</caption><thead><tr><th></th><th>PLA</th><th>Steel C40</th></tr></thead><tbody><tr><td>Density (ρ) [g/cm³]</td><td>1,21-1,25</td><td>7,85</td></tr><tr><td>Tensile strength (σ) [MPa]</td><td>21-60</td><td>650-800</td></tr><tr><td>Young's modulus (E) [GPa]</td><td>0,35-3,5</td><td>210</td></tr><tr><td>Elongation (ε) [%]</td><td>2,5-6</td><td>20-25</td></tr><tr><td>Glass transition temperature (Tg) [°C]</td><td>45-60</td><td>-</td></tr><tr><td>Melting temperature (Tm) [°C]</td><td>150-162</td><td>-</td></tr></tbody></table>

Fig. 2 - displays the geometries of the metal and polymeric partners, based on the ASTM: D5961 standard, which provides details on the execution of the lap shear test (Fig. 2 - b). The steel C40 sheet has been drilled according to the suggested hole size, while the PLA sheet has been produced by additive manufacturing and it was characterized by a rivet pin height of 8 mm.
<figure><img src="image_1600.png" /><figcaption>Fig. 2 - Staking process principle</figcaption></figure>

As previously introduced, the FDM, a versatile and cost-effective 3D technique, allowed to create the thermoplastic component layer by layer. More in detail, printer heats the thermoplastic filament (ø = 1.75 mm) which is fed in the extrusion nozzle and deposited layer by layer on a build platform to create the desired solid objects based on a digital and sliced model [13]. To the aim of this study, two different printers were used in order to test their influence, working at the printing speed recommended by the Bambu Studio software, which was used for slicing, according to the selected printer. The printing speed is the velocity at which the nozzle moves in relation to the printing platform.

Sample 1 was produced by the Creality Ender-3 V2 (Fig. 3 – a), characterized by a recommended print speed of 50 mm/s, while Sample 2 by the Bambu Lab P1P (Fig. 3 – b) characterized by a print speed of 270 mm/s.

<figure><img src="image_1601.png" /><figcaption>Fig. 3 – a) Creality Ender-3 V2 and b) Bambu Lab P1P</figcaption></figure>

The overlapped polymeric and the metallic plates (Fig. 4 – a) were joined using the Hot-Air Cold Staking (HACS) technique. The pin rivet was heated by a hot air gun and the temperature was monitored by a thermocouple inserted in the middle of the pin using a drilled hole on the opposite face of the sample respect to the hot air jet. As soon as a temperature of 55°C was reached, the pin was shaped using a cold deformation tool, installed on a manual hydraulic press, whose movement was fixed and maintained constant for the all tests (Fig. 4 – b). Flat-headed rivets, characterised by the same height of the head, were obtained creating the joints between the polymeric and metallic plates.
<figure><img src="image_1601.png" /><figcaption>Fig. 2 - Staking process principle</figcaption></figure>

As previously introduced, the FDM, a versatile and cost-effective 3D technique, allowed to create the thermoplastic component layer by layer. More in detail, printer heats the thermoplastic filament (ø = 1.75 mm) which is fed in the extrusion nozzle and deposited layer by layer on a build platform to create the desired solid objects based on a digital and sliced model [13]. To the aim of this study, two different printers were used in order to test their influence, working at the printing speed recommended by the Bambu Studio software, which was used for slicing, according to the selected printer. The printing speed is the velocity at which the nozzle moves in relation to the printing platform.

Sample 1 was produced by the Creality Ender-3 V2 (Fig. 3 – a), characterized by a recommended print speed of 50 mm/s, while Sample 2 by the Bambu Lab P1P (Fig. 3 – b) characterized by a print speed of 270 mm/s.

<figure><img src="image_1602.png" /><figcaption>Fig. 3 – a) Creality Ender-3 V2 and b) Bambu Lab P1P</figcaption></figure>

The overlapped polymeric and the metallic plates (Fig. 4 – a) were joined using the Hot-Air Cold Staking (HACS) technique. The pin rivet was heated by a hot air gun and the temperature was monitored by a thermocouple inserted in the middle of the pin using a drilled hole on the opposite face of the sample respect to the hot air jet. As soon as a temperature of 55°C was reached, the pin was shaped using a cold deformation tool, installed on a manual hydraulic press, whose movement was fixed and maintained constant for the all tests (Fig. 4 – b). Flat-headed rivets, characterised by the same height of the head, were obtained creating the joints between the polymeric and metallic plates.
<figure><img src="image_1602.png" /><figcaption>Fig. 4 – Experimental test of Hot-Air Cold Staking: a) partners positioning, b) joint execution</figcaption></figure>

The details of the process for printing the polymeric samples are schematised in the following Table 2.

<table><caption>Table 2 – Experimental details of the polymeric parts</caption><thead><tr><td>Sample</td><td>Printer</td><td>Print speed [mm/s]</td><td>Total printing time [min]</td><td>Filament used [g]</td></tr></thead><tbody><tr><td>1</td><td>Creality Ender-3 V2</td><td>50</td><td>128,70</td><td>26,54</td></tr><tr><td>2</td><td>Bambu Lab P1P</td><td>270</td><td>30,28</td><td>26,98</td></tr></tbody></table>

Metal and polymeric components were prepared and joined, as discussed above (Fig. 5 – a, b). Lap shear tests were performed using the MTS Criterion Series 40 testing machine (Fig. 5 – c). Considering the ASTM: D5961 standard procedure, a traverse speed of 2 mm/min was set. Three replicates for each processing condition were investigated at room temperature and curves for each test were obtained recording load and displacement.

<figure><img src="image_1603.png" /><figcaption>Fig. 5 – a) Metallic and polymeric partners, b) hybrid joint (Test 1), c) lap shear test</figcaption></figure>
<figure><img src="image_1603.png" /><figcaption>Fig. 4 – Experimental test of Hot-Air Cold Staking: a) partners positioning, b) joint execution</figcaption></figure>

The details of the process for printing the polymeric samples are schematised in the following Table 2.

<table><caption>Table 2 – Experimental details of the polymeric parts</caption><thead><tr><td>Sample</td><td>Printer</td><td>Print speed [mm/s]</td><td>Total printing time [min]</td><td>Filament used [g]</td></tr></thead><tbody><tr><td>1</td><td>Creality Ender-3 V2</td><td>50</td><td>128,70</td><td>26,54</td></tr><tr><td>2</td><td>Bambu Lab P1P</td><td>270</td><td>30,28</td><td>26,98</td></tr></tbody></table>

Metal and polymeric components were prepared and joined, as discussed above (Fig. 5 – a, b). Lap shear tests were performed using the MTS Criterion Series 40 testing machine (Fig. 5 – c). Considering the ASTM: D5961 standard procedure, a traverse speed of 2 mm/min was set. Three replicates for each processing condition were investigated at room temperature and curves for each test were obtained recording load and displacement.

<figure><img src="image_1604.png" /><figcaption>Fig. 5 – a) Metallic and polymeric partners, b) hybrid joint (Test 1), c) lap shear test</figcaption></figure>
## Results and discussion

The load-displacement curves in Fig. 6 illustrate the effect of the type of the printing machine on the strength of the joints. Test 1, produced using Sample 1, which was printed by means of the slowest FDM machine (Fig. 3 - a), exhibits higher mechanical performance in the all replicates. This result is supported by a higher maximum load and the ability to withstand greater displacement before reaching the breaking point. On the other hand, lower strength is observable in case of Sample 2, where the increasing printing speed seems compromising the interface bond strengths between the deposited layers within the polymeric partners. The reason may be related to the plasticization time of the polymer. In fact, a higher printing speed does not allow the material to plasticize in a proper manner, increasing the residual stress produced and therefore weaker mechanical properties of the final hybrid joint are observed. The second replicate of Sample 2 shows a special behaviour, which would deserve an in-depth analysis.

<figure><img src="image_1604.png" /><figcaption>Fig. 6 - Load-displacement curves for all replicates of Sample 1 (print speed=50 mm/s) and Sample 2 (print speed=270 mm/s)</figcaption></figure>

The aforementioned considerations were also supported by the experimental evidences. In fact, Fig. 7 displays the samples at the end of the lap shear tests. Fig. 7 - a refers to Sample 1 and the pin rivet experienced a slower detachment, while the fracture occurred subsequently at the base of the pinhead after detaching from the metal partner. In contrast, in Sample 2 (Fig. 7 - b), produced at higher speed, the poor material plasticization appears evident also here from the detachment of the layers from the base.

<figure><img src="image_1605.png" /><figcaption>Fig. 7 - Failure mode of a) Sample 1 and b) Sample 2</figcaption></figure>
## Results and discussion

The load-displacement curves in Fig. 6 illustrate the effect of the type of the printing machine on the strength of the joints. Test 1, produced using Sample 1, which was printed by means of the slowest FDM machine (Fig. 3 - a), exhibits higher mechanical performance in the all replicates. This result is supported by a higher maximum load and the ability to withstand greater displacement before reaching the breaking point. On the other hand, lower strength is observable in case of Sample 2, where the increasing printing speed seems compromising the interface bond strengths between the deposited layers within the polymeric partners. The reason may be related to the plasticization time of the polymer. In fact, a higher printing speed does not allow the material to plasticize in a proper manner, increasing the residual stress produced and therefore weaker mechanical properties of the final hybrid joint are observed. The second replicate of Sample 2 shows a special behaviour, which would deserve an in-depth analysis.

<figure><img src="image_1605.png" /><figcaption>Fig. 6 - Load-displacement curves for all replicates of Sample 1 (print speed=50 mm/s) and Sample 2 (print speed=270 mm/s)</figcaption></figure>

The aforementioned considerations were also supported by the experimental evidences. In fact, Fig. 7 displays the samples at the end of the lap shear tests. Fig. 7 - a refers to Sample 1 and the pin rivet experienced a slower detachment, while the fracture occurred subsequently at the base of the pinhead after detaching from the metal partner. In contrast, in Sample 2 (Fig. 7 - b), produced at higher speed, the poor material plasticization appears evident also here from the detachment of the layers from the base.

<figure><img src="image_1606.png" /><figcaption>Fig. 7 - Failure mode of a) Sample 1 and b) Sample 2</figcaption></figure>
## Conclusions

In this work, hybrid joints were experimentally investigated. The polymeric partners were produced by additive manufacturing, specifically by FDM, using PLA as base material and joined with a holed steel plate. The joint was performed by the Hot-Air Cold Staking (HACS) technique. The polymeric partners were produced by means of two different printers, which allowed the setting of two printing speeds. The influence of these latter on the final joint strength was evaluated by lap shear test which revealed that a slower printing speed (50 mm/s) yielded a stronger joint, attributed to enhanced material plasticization and the consequent formation of robust bonds. Conversely, the machine operating at a higher printing speed (270 mm/s) demonstrated increased process efficiency but compromised joint strength. A distinctive failure mode was also observed at the end of the lap shear test. The delamination of layers from the polymeric base emphasized the impact of reduced material plasticization at higher speeds. These preliminary results allowed to reflect on the joining process planning based on the characteristics that the produced pieces have to possess. This involves also the evaluation of the optimization of the process duration according to the structural requirements of the parts.

Future experimental tests will be performed in order to investigate further printing speeds on the same printers to take into account the machine stiffness and the related deposition load that could affect the intimate contact for better diffusion of polymer chains across the interface at subsequent layers resulting in the formation of strong bonds. In addition, experimental results will be verified by numerical investigations, aiming at proving a tool useful for optimizing the staking process for joining polymers and metals according to the specific context where thy hybrid joints could be used.

## References

[1] K. Martinsen, S.J. Hu, B.E. Carlson, Joining of dissimilar materials, CIRP Ann., 64 (2015) 679-699. https://doi.org/10.1016/j.cirp.2015.05.006

[2] A.B. Abibe, S.T. Amancio-Filho, J.F. dos Santos, E. Hage, Mechanical and failure behaviour of hybrid polymer-metal staked joints, Mater. Des., 46 (2013) 338-347. https://doi.org/10.1016/j.matdes.2012.10.043

[3] F.C. Liu, K. Nakata, J. Liao, S. Hirota, H. Fukui, Reducing bubbles in friction lap welded joint of magnesium alloy and polyamide, Sci. Technol. Weld. Join., 19 (2014) 578-587. https://doi.org/10.1179/1362171814Y.0000000228

[4] F.C. Liu, P. Dong, W. Lu, K. Sun, On formation of Al O C bonds at aluminum/polyamide joint interface, Appl. Surf. Sci., 466 (2019) 202-209. https://doi.org/10.1016/j.apsusc.2018.10.024

[5] A.B. Abibe, M. Sônego, J.F. dos Santos, L.B. Canto, S.T. Amancio-Filho, On the feasibility of a friction-based staking joining method for polymer-metal hybrid structures, Mater. Des., 92 (2016) 632-642. https://doi.org/10.1016/j.matdes.2015.12.087

[6] Y. Huang, X. Meng, Y. Xie, J. Li, L. Wan, New technique of friction-based filling stacking joining for metal and polymer, Compos. Part B Eng., 163 (2019) 217-223. https://doi.org/10.1016/j.compositesb.2018.11.050

[7] F.C. Liu, P. Dong, X. Pei, A high-speed metal-to-polymer direct joining technique and underlying bonding mechanisms, J. Mater. Process. Technol., 280 (2020) 116610. https://doi.org/10.1016/j.jmatprotec.2020.116610

[8] Y. Huang, X. Gao, Y. Zhang, B. Ma, Laser joining technology of polymer-metal hybrid structures - A review, J. Manuf. Process., 79 (2022) 934-961. https://doi.org/10.1016/j.jmapro.2022.05.026
## Conclusions

In this work, hybrid joints were experimentally investigated. The polymeric partners were produced by additive manufacturing, specifically by FDM, using PLA as base material and joined with a holed steel plate. The joint was performed by the Hot-Air Cold Staking (HACS) technique. The polymeric partners were produced by means of two different printers, which allowed the setting of two printing speeds. The influence of these latter on the final joint strength was evaluated by lap shear test which revealed that a slower printing speed (50 mm/s) yielded a stronger joint, attributed to enhanced material plasticization and the consequent formation of robust bonds. Conversely, the machine operating at a higher printing speed (270 mm/s) demonstrated increased process efficiency but compromised joint strength. A distinctive failure mode was also observed at the end of the lap shear test. The delamination of layers from the polymeric base emphasized the impact of reduced material plasticization at higher speeds. These preliminary results allowed to reflect on the joining process planning based on the characteristics that the produced pieces have to possess. This involves also the evaluation of the optimization of the process duration according to the structural requirements of the parts.

Future experimental tests will be performed in order to investigate further printing speeds on the same printers to take into account the machine stiffness and the related deposition load that could affect the intimate contact for better diffusion of polymer chains across the interface at subsequent layers resulting in the formation of strong bonds. In addition, experimental results will be verified by numerical investigations, aiming at proving a tool useful for optimizing the staking process for joining polymers and metals according to the specific context where thy hybrid joints could be used.

## References

[1] K. Martinsen, S.J. Hu, B.E. Carlson, Joining of dissimilar materials, CIRP Ann., 64 (2015) 679-699. https://doi.org/10.1016/j.cirp.2015.05.006

[2] A.B. Abibe, S.T. Amancio-Filho, J.F. dos Santos, E. Hage, Mechanical and failure behaviour of hybrid polymer-metal staked joints, Mater. Des., 46 (2013) 338-347. https://doi.org/10.1016/j.matdes.2012.10.043

[3] F.C. Liu, K. Nakata, J. Liao, S. Hirota, H. Fukui, Reducing bubbles in friction lap welded joint of magnesium alloy and polyamide, Sci. Technol. Weld. Join., 19 (2014) 578-587. https://doi.org/10.1179/1362171814Y.0000000228

[4] F.C. Liu, P. Dong, W. Lu, K. Sun, On formation of Al O C bonds at aluminum/polyamide joint interface, Appl. Surf. Sci., 466 (2019) 202-209. https://doi.org/10.1016/j.apsusc.2018.10.024

[5] A.B. Abibe, M. Sônego, J.F. dos Santos, L.B. Canto, S.T. Amancio-Filho, On the feasibility of a friction-based staking joining method for polymer-metal hybrid structures, Mater. Des., 92 (2016) 632-642. https://doi.org/10.1016/j.matdes.2015.12.087

[6] Y. Huang, X. Meng, Y. Xie, J. Li, L. Wan, New technique of friction-based filling stacking joining for metal and polymer, Compos. Part B Eng., 163 (2019) 217-223. https://doi.org/10.1016/j.compositesb.2018.11.050

[7] F.C. Liu, P. Dong, X. Pei, A high-speed metal-to-polymer direct joining technique and underlying bonding mechanisms, J. Mater. Process. Technol., 280 (2020) 116610. https://doi.org/10.1016/j.jmatprotec.2020.116610

[8] Y. Huang, X. Gao, Y. Zhang, B. Ma, Laser joining technology of polymer-metal hybrid structures - A review, J. Manuf. Process., 79 (2022) 934-961. https://doi.org/10.1016/j.jmapro.2022.05.026
[9] S.T. Amancio-Filho, L.-A. Blaga, Joining of Polymer–Metal Hybrid Structures Principles and Applications, 2018. https://doi.org/10.1002/9781119429807

[10] S. Härtel, E. Brueckner, B. Awiszus, M. Gehde, Development of a Numerical Model of the Hot Air Staking Process Based on Experimental Data, Appl. Sci., 10 (2020) 7115. https://doi.org/10.3390/app10207115

[11] S. Farah, D.G. Anderson, R. Langer, Physical and mechanical properties of PLA, and their functions in widespread applications — A comprehensive review, Adv. Drug Deliv. Rev., 107 (2016) 367–392. https://doi.org/10.1016/j.amdr.2016.06.012

[12] L. Sandanamsamy, J. Mogan, K. Rajan, W.S.W. Harun, I. Ishak, F.R.M. Romlay, M. Samykano, K. Kadirgama, Effect of process parameter on tensile properties of FDM printed PLA, Mater. Today Proc., (2023). https://doi.org/10.1016/j.matpr.2023.03.217

[13] R. Patel, C. Desai, S. Kushwah, M.H. Mangrola, A review article on FDM process parameters in 3D printing for composite materials, Mater. Today Proc., 60 (2022) 2162–2166. https://doi.org/10.1016/j.matpr.2022.02.385
[9] S.T. Amancio-Filho, L.-A. Blaga, Joining of Polymer–Metal Hybrid Structures Principles and Applications, 2018. https://doi.org/10.1002/9781119429807

[10] S. Härtel, E. Brueckner, B. Awiszus, M. Gehde, Development of a Numerical Model of the Hot Air Staking Process Based on Experimental Data, Appl. Sci., 10 (2020) 7115. https://doi.org/10.3390/app10207115

[11] S. Farah, D.G. Anderson, R. Langer, Physical and mechanical properties of PLA, and their functions in widespread applications — A comprehensive review, Adv. Drug Deliv. Rev., 107 (2016) 367–392. https://doi.org/10.1016/j.amdr.2016.06.012

[12] L. Sandanamsamy, J. Mogan, K. Rajan, W.S.W. Harun, I. Ishak, F.R.M. Romlay, M. Samykano, K. Kadirgama, Effect of process parameter on tensile properties of FDM printed PLA, Mater. Today Proc., (2023). https://doi.org/10.1016/j.matpr.2023.03.217

[13] R. Patel, C. Desai, S. Kushwah, M.H. Mangrola, A review article on FDM process parameters in 3D printing for composite materials, Mater. Today Proc., 60 (2022) 2162–2166. https://doi.org/10.1016/j.matpr.2022.02.385
# Numerical investigation of rotational friction welding for C22.8 - 41Cr4 joints using a substitute model

MOHNFELD Norman¹,*, WESTER Hendrik¹, TUNC KARAER Gökhan¹, PIWEK Armin¹ and UHE Johanna¹,ᵇ

¹Institute of Forming Technology and Machines (LUH), An der Universität 2, 30823 Garbsen, Germany

ªmohnfeld@ifum.uni-hannover.de, ᵇuhe@ifum.uni-hannover.de

**Keywords:** Rotational Friction Welding, FEM, Hybrid Components, Substitute Model

**Abstract.** Rotational friction welding (RFW) is a solid-state joining process that enables the joining of similar and dissimilar materials such as metal-metal or metal-ceramic joints. Due to its high application flexibility, this process has great potential for the production of hybrid components. In order to be able to realise this potential for the production of hybrid components, the development of an improved process design is required. Due to the complexity of the process, the Finite Element Method (FEM) can be used to solve complex problems and is therefore an established tool for the design of joining processes. This work focuses on the development of an FE model to represent the RFW process of C22.8 and 41Cr4 joints. The material data required for the numerical representation of the RFW were obtained from isothermal cylinder compression tests. The frictional heat which is generated during RFW is calculated using a substitute model, which mainly depends on the Y-factor. The Y-factor indicates what percentage of the calculated frictional energy is introduced into the process. The Y-factor was determined and then verified using experimental data. A general validity of the determined Y-factors with changed process parameters could not be achieved.

**Introduction**

Rotational friction welding (RFW) as a solid-state joining process represents a promising process for the economic production of hybrid high-performance components. Compared to conventional fusion-metallurgical welding processes, the RFW is characterised by short welding times, high geometrical accuracy, low energy consumption, excellent static and dynamic bond strengths with fine grain structure and large material variety [1]. Furthermore, RFW can be used to join materials that can only be welded to a limited extent or not at all using conventional fusion welding processes (e.g. aluminium-copper or steel-ceramic) [2]. The design of RFW processes is complicated due to the rapid process dynamics and the complexity of the material phenomena involved. With regard to the methodical design and development of production processes, the finite element method (FEM) is a proven instrument for tool and process design of bulk metal forming as well as joining processes. Time-consuming and cost-intensive steps in the product development phase can be reduced and the existing process limits can be systematically expanded through virtual testing of new materials or process variants [3]. Soucail et al. investigated a thermo-mechanically coupled FE simulation of friction welding process using the lagrangian formulation with an explicit scheme. A good agreement with the experimentally measured temperature profiles near the weld seam could be achieved. At a distance of 0.3 mm, however, the calculation results deviated considerably from the experimental measurements [4]. A cooling simulation decoupled from the friction welding process was carried out by Nguyn et al.. Cooling curves were taken from the experiment and applied to the workpieces. Using microstructure models and experimental time-temperature-conversion diagrams, an estimation of the phase fractions after cooling could be made, whereby the workpieces were modelled in the FE simulation as rigid bodies without plastic

<figure><img src="image_1606.png" /></figure>
# Numerical investigation of rotational friction welding for C22.8 - 41Cr4 joints using a substitute model

MOHNFELD Norman¹,*, WESTER Hendrik¹, TUNC KARAER Gökhan¹, PIWEK Armin¹ and UHE Johanna¹,ᵇ

¹Institute of Forming Technology and Machines (LUH), An der Universität 2, 30823 Garbsen, Germany

a>mohnfeld@ifum.uni-hannover.de, b>uhe@ifum.uni-hannover.de

**Keywords:** Rotational Friction Welding, FEM, Hybrid Components, Substitute Model

**Abstract.** Rotational friction welding (RFW) is a solid-state joining process that enables the joining of similar and dissimilar materials such as metal-metal or metal-ceramic joints. Due to its high application flexibility, this process has great potential for the production of hybrid components. In order to be able to realise this potential for the production of hybrid components, the development of an improved process design is required. Due to the complexity of the process, the Finite Element Method (FEM) can be used to solve complex problems and is therefore an established tool for the design of joining processes. This work focuses on the development of an FE model to represent the RFW process of C22.8 and 41Cr4 joints. The material data required for the numerical representation of the RFW were obtained from isothermal cylinder compression tests. The frictional heat which is generated during RFW is calculated using a substitute model, which mainly depends on the Y-factor. The Y-factor indicates what percentage of the calculated frictional energy is introduced into the process. The Y-factor was determined and then verified using experimental data. A general validity of the determined Y-factors with changed process parameters could not be achieved.

**Introduction**

Rotational friction welding (RFW) as a solid-state joining process represents a promising process for the economic production of hybrid high-performance components. Compared to conventional fusion-metallurgical welding processes, the RFW is characterised by short welding times, high geometrical accuracy, low energy consumption, excellent static and dynamic bond strengths with fine grain structure and large material variety [1]. Furthermore, RFW can be used to join materials that can only be welded to a limited extent or not at all using conventional fusion welding processes (e.g. aluminium-copper or steel-ceramic) [2]. The design of RFW processes is complicated due to the rapid process dynamics and the complexity of the material phenomena involved. With regard to the methodical design and development of production processes, the finite element method (FEM) is a proven instrument for tool and process design of bulk metal forming as well as joining processes. Time-consuming and cost-intensive steps in the product development phase can be reduced and the existing process limits can be systematically expanded through virtual testing of new materials or process variants [3]. Soucail et al. investigated a thermo-mechanically coupled FE simulation of friction welding process using the lagrangian formulation with an explicit scheme. A good agreement with the experimentally measured temperature profiles near the weld seam could be achieved. At a distance of 0.3 mm, however, the calculation results deviated considerably from the experimental measurements [4]. A cooling simulation decoupled from the friction welding process was carried out by Nguyn et al.. Cooling curves were taken from the experiment and applied to the workpieces. Using microstructure models and experimental time-temperature-conversion diagrams, an estimation of the phase fractions after cooling could be made, whereby the workpieces were modelled in the FE simulation as rigid bodies without plastic

<figure><img src="image_1607.png" /></figure>

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deformation [5]. Schmicker conducted a simulation study of the friction welding of a mono material steel joint. Using fluid-dynamic material models, a high reproduction quality of the plastic material flow could be achieved in a MATLAB environment. However, due to the lack of structural-mechanical coupling, the elastic strains and thermal shrinkage were not taken into account [6]. Nu et al. numerically investigated a friction welded joint made of a Ti-alloy. The frictional heat was simplified as a constant heat flux density, which was interrupted when the melting temperature was reached. Although experimental investigations were presented in the paper, no comparison was made with the numerical results [7]. A welded joint made of a Ni alloy was also modelled by Wang et al. using ABAQUS. In the process parameter study, the influence of the rotational speed on the resulting part shortening and the temperature development was investigated without an experimental validation of the results [8]. In a study by Li et al. the temperature development at the beginning of the process was modelled with analytical approaches. Validation was based on a peak temperature measurement using a thermographic camera [9]. Friction welding processes for a Ni-alloy and a CrNi stainless steel were modelled with DEFORM by Nan et al. In principle, a good temperature prediction could be achieved, considering only a locally undefined peak temperature at the surface. From the comparison of the weld flash geometry, it could be deduced that the component geometry was underestimated by the FE model [10].

In summary, either the temperature development was considered without including the plastic deformation or the forming was simulated and the heating was modelled by an external heat source.

Using RFW, similar and dissimilar joints can be produced that have sufficient strength for subsequent forming. RFW is therefore a promising approach for the production of hybrid components. In the Collaborative Research Centre (CRC) 1153 Tailored Forming, hybrid semi-finished products consisting of different steels or steel and aluminium are produced using RFW [11]. The design of the RFW processes as well as the investigation of the influence of the specimen geometries on the composite strengths was determined by experimental tests in previous studies of the authors [12,13]. By a numerical consideration of the RFW process, extensive experimental investigations can be reduced on the one hand and on the other hand the analyses in the joining zone can be carried out without destroying the component. Furthermore, it is possible to integrate subroutines into the model so that, for example, the microstructure transformation processes or the bond strength can be represented.

This work deals with the numerical representation of the RFW process of C22.8 and 41Cr4. In order to describe the thermomechanical material behaviour during friction welding, process flow properties are determined by means of isothermal cylinder compression tests in a temperature range from 20 °C to 1,400 °C. Based on the results analytical material models are created and implemented in the FE-software using subroutines. The RFW process is setup as a 2D model in the commercial FE software Simufact forming v16 (simufact engineering GmbH, Hamburg, Germany). First, the influence of the mesh size on the results is analysed. Subsequently, the Y-factor of the substitute model for calculating the frictional heat in the process is examined and investigated. Finally, the simulation model is verified using the inversely determined Y-factors and finally validated by varying the various process parameters.

## Methods - Experimental Investigation

**Experimental Investigation:** To determine the chemical composition of the materials used, measurements were carried out on the SPECTROMAXx radio spectrometer in argon atmosphere. Each material was radioed nine times for statistical validation of the measurement. The results of the radio spectroscopy measurement of the investigated materials are listed in Table 1. It is noticeable that for both materials a carbon value slightly below the prescribed minimum value was measured. In addition, 41Cr4 has a slightly higher chromium content. All other measured values are within the specified tolerance ranges.
deformation [5]. Schmicker conducted a simulation study of the friction welding of a mono material steel joint. Using fluid-dynamic material models, a high reproduction quality of the plastic material flow could be achieved in a MATLAB environment. However, due to the lack of structural-mechanical coupling, the elastic strains and thermal shrinkage were not taken into account [6]. Nu et al. numerically investigated a friction welded joint made of a Ti-alloy. The frictional heat was simplified as a constant heat flux density, which was interrupted when the melting temperature was reached. Although experimental investigations were presented in the paper, no comparison was made with the numerical results [7]. A welded joint made of a Ni alloy was also modelled by Wang et al. using ABAQUS. In the process parameter study, the influence of the rotational speed on the resulting part shortening and the temperature development was investigated without an experimental validation of the results [8]. In a study by Li et al. the temperature development at the beginning of the process was modelled with analytical approaches. Validation was based on a peak temperature measurement using a thermographic camera [9]. Friction welding processes for a Ni-alloy and a CrNi stainless steel were modelled with DEFORM by Nan et al. In principle, a good temperature prediction could be achieved, considering only a locally undefined peak temperature at the surface. From the comparison of the weld flash geometry, it could be deduced that the component geometry was underestimated by the FE model [10].

In summary, either the temperature development was considered without including the plastic deformation or the forming was simulated and the heating was modelled by an external heat source.

Using RFW, similar and dissimilar joints can be produced that have sufficient strength for subsequent forming. RFW is therefore a promising approach for the production of hybrid components. In the Collaborative Research Centre (CRC) 1153 Tailored Forming, hybrid semi-finished products consisting of different steels or steel and aluminium are produced using RFW [11]. The design of the RFW processes as well as the investigation of the influence of the specimen geometries on the composite strengths was determined by experimental tests in previous studies of the authors [12,13]. By a numerical consideration of the RFW process, extensive experimental investigations can be reduced on the one hand and on the other hand the analyses in the joining zone can be carried out without destroying the component. Furthermore, it is possible to integrate subroutines into the model so that, for example, the microstructure transformation processes or the bond strength can be represented.

This work deals with the numerical representation of the RFW process of C22.8 and 41Cr4. In order to describe the thermomechanical material behaviour during friction welding, process flow properties are determined by means of isothermal cylinder compression tests in a temperature range from 20 °C to 1,400 °C. Based on the results analytical material models are created and implemented in the FE-software using subroutines. The RFW process is setup as a 2D model in the commercial FE software Simufact forming v16 (simufact engineering GmbH, Hamburg, Germany). First, the influence of the mesh size on the results is analysed. Subsequently, the Y-factor of the substitute model for calculating the frictional heat in the process is examined and investigated. Finally, the simulation model is verified using the inversely determined Y-factors and finally validated by varying the various process parameters.

## Methods - Experimental Investigation

Experimental Investigation: To determine the chemical composition of the materials used, measurements were carried out on the SPECTROMAXx radio spectrometer in argon atmosphere. Each material was radioed nine times for statistical validation of the measurement. The results of the radio spectroscopy measurement of the investigated materials are listed in Table 1. It is noticeable that for both materials a carbon value slightly below the prescribed minimum value was measured. In addition, 41Cr4 has a slightly higher chromium content. All other measured values are within the specified tolerance ranges.
<table><caption>Table 1. Chemical composition of the C22.8 and 41Cr4 steels in wt.%, measured by optical emission spectrometry.</caption><thead><tr><th>Element</th><th></th><th>C</th><th>Si</th><th>Mn</th><th>P</th><th>S</th><th>Cr</th><th>Mo</th><th>Ni</th><th>Al</th><th>Co</th><th>Cu</th></tr></thead><tbody><tr><td rowspan="3">C22.8</td><td>Max</td><td>0.23</td><td>0.4</td><td>0.9</td><td>0.025</td><td>0.015</td><td>0.3</td><td>0.1</td><td>0.3</td><td>0.05</td><td>0.1</td><td>0.4</td></tr><tr><td>Min</td><td>0.18</td><td>-</td><td>0.3</td><td>-</td><td>-</td><td>-</td><td>-</td><td>-</td><td>0.015</td><td>-</td><td>-</td></tr><tr><td>Test</td><td><strong>0.166</strong></td><td><strong>0.208</strong></td><td><strong>0.817</strong></td><td><strong>0.005</strong></td><td><strong>0.005</strong></td><td><strong>0.116</strong></td><td><strong>0.022</strong></td><td><strong>0.095</strong></td><td><strong>0.023</strong></td><td><strong>0.008</strong></td><td><strong>0.196</strong></td></tr><tr><td rowspan="3">41Cr4</td><td>Max</td><td>0.45</td><td>0.4</td><td>0.9</td><td>0.035</td><td>0.03</td><td>1.2</td><td>0.1</td><td>0.3</td><td>-</td><td>0.1</td><td>0.4</td></tr><tr><td>Min</td><td>0.38</td><td>-</td><td>0.6</td><td>-</td><td>-</td><td>0.9</td><td>-</td><td>-</td><td>-</td><td>-</td><td>-</td></tr><tr><td>Test</td><td><strong>0.36</strong></td><td><strong>0.305</strong></td><td><strong>0.795</strong></td><td><strong>0.004</strong></td><td><strong>0.017</strong></td><td><strong>1.24</strong></td><td><strong>0.01</strong></td><td><strong>0.024</strong></td><td><strong>0.001</strong></td><td><strong>0.001</strong></td><td><strong>0.013</strong></td></tr></tbody></table>

To describe the flow behaviour cylinder compression tests were carried out for C22.8 and 41Cr4 according to the parameters listed in Table 2.

<table><caption>Table 2. Experimental parameters for the cylinder compression tests of C22.8 and 41Cr4.</caption><thead><tr><th>material</th><th>strain rate in s<sup>-1</sup></th><th>temperature in °C</th></tr></thead><tbody><tr><td>C22.8</td><td>0.1; 1; 10</td><td>1,300 and 1,400</td></tr><tr><td>41Cr4</td><td>0.1; 1; 10</td><td>1,300 and 1,400</td></tr></tbody></table>

The specimen geometry was Ø 10 mm x 15 mm and specimens were prepared from plane-parallel discs by wire erosion. The compression tests were carried out on the Gleeble 3800 GTC forming simulator. Under vacuum, the specimens were heated conductively at a heating rate of 4.5 K/s and held at the target temperature for 30 s to realise a homogeneous temperature distribution. Since partially liquid phases are already formed at this temperature, the specimens are compressed in a quartz sleeve. The quartz sleeve shields the specimen so that no particles can damage the test chamber. The punches are made of solid hard metal, which has a very high heat resistance. For the strain rates of 1 s⁻¹ and 10 s⁻¹ the tests were displacement controlled and for 0.1 s⁻¹ strain controlled. The specimens were upset to a plastic strain of 0.7. Each parameter combination from Table 2 was repeated five times. For the temperature range below 1,300 °C up to room temperature, flow curves were also recorded from the same material batch by cylinder compression tests [14].

The experimental investigations of the RFW process were carried out on the Genius Plus RS30 from KUKA. On the spindle side, a cylinder made of 41Cr4 with a Ø 40 mm and a length of 160 mm was inserted. A cylinder made of C22.8 with a Ø 40 mm and a length of 210 mm was used on the sledge side. The temperature measurement was carried out with a type S thermocouple. According to Fig. 1 (a), eight type K thermocouples (TC) were attached to the C22.8 bolts. In order to measure the temperature in the contact surface of the friction partners TC1 and TC5 were guided directly into the end face via small grooves. At 90° to this, measurements were taken axially around the circumference at a distance of 5 mm, 10 mm and 15 mm from the end face (TC2-TC4 and TC6-TC8). This allows the heat dissipation along the specimen to be determined. In addition, force, displacement as well as the rotational speed are recorded during the experiments. The applied time-rotational speed and force-time profiles for the 1 parameter set and 2 parameter set is shown in Fig. 1 (b). In the friction phase, heat is generated by the rotation and the applied axial force. By stopping the spindle, the upsetting phase is initiated and the bond is created. Fig. 1 (a) shows the process at the beginning of the friction phase on the left and after the end of the upsetting phase on the right. The RFW tests were repeated five times using this 1 parameter set with a rotational speed of 1,500 min⁻¹ and a upsetting force of 180 kN. In addition, a test with increased rotational speed of 2,000 min⁻¹ and increased upsetting force of 220 kN was carried out for later model validation in 2 parameter set. One specimen was shortened and separated along the
<table><caption>Table 1. Chemical composition of the C22.8 and 41Cr4 steels in wt.%, measured by optical emission spectrometry.</caption><thead><tr><th>Element</th><th></th><th>C</th><th>Si</th><th>Mn</th><th>P</th><th>S</th><th>Cr</th><th>Mo</th><th>Ni</th><th>Al</th><th>Co</th><th>Cu</th></tr></thead><tbody><tr><td rowspan="3">C22.8</td><td>Max</td><td>0.23</td><td>0.4</td><td>0.9</td><td>0.025</td><td>0.015</td><td>0.3</td><td>0.1</td><td>0.3</td><td>0.05</td><td>0.1</td><td>0.4</td></tr><tr><td>Min</td><td>0.18</td><td>-</td><td>0.3</td><td>-</td><td>-</td><td>-</td><td>-</td><td>-</td><td>0.015</td><td>-</td><td>-</td></tr><tr><td>Test</td><td><strong>0.166</strong></td><td><strong>0.208</strong></td><td><strong>0.817</strong></td><td><strong>0.005</strong></td><td><strong>0.005</strong></td><td><strong>0.116</strong></td><td><strong>0.022</strong></td><td><strong>0.095</strong></td><td><strong>0.023</strong></td><td><strong>0.008</strong></td><td><strong>0.196</strong></td></tr><tr><td rowspan="3">41Cr4</td><td>Max</td><td>0.45</td><td>0.4</td><td>0.9</td><td>0.035</td><td>0.03</td><td>1.2</td><td>0.1</td><td>0.3</td><td>-</td><td>0.1</td><td>0.4</td></tr><tr><td>Min</td><td>0.38</td><td>-</td><td>0.6</td><td>-</td><td>-</td><td>0.9</td><td>-</td><td>-</td><td>-</td><td>-</td><td>-</td></tr><tr><td>Test</td><td><strong>0.36</strong></td><td><strong>0.305</strong></td><td><strong>0.795</strong></td><td><strong>0.004</strong></td><td><strong>0.017</strong></td><td><strong>1.24</strong></td><td><strong>0.01</strong></td><td><strong>0.024</strong></td><td><strong>0.001</strong></td><td><strong>0.001</strong></td><td><strong>0.013</strong></td></tr></tbody></table>

To describe the flow behaviour cylinder compression tests were carried out for C22.8 and 41Cr4 according to the parameters listed in Table 2.

<table><caption>Table 2. Experimental parameters for the cylinder compression tests of C22.8 and 41Cr4.</caption><thead><tr><th>material</th><th>strain rate in s<sup>-1</sup></th><th>temperature in °C</th></tr></thead><tbody><tr><td>C22.8</td><td>0.1; 1; 10</td><td>1,300 and 1,400</td></tr><tr><td>41Cr4</td><td>0.1; 1; 10</td><td>1,300 and 1,400</td></tr></tbody></table>

The specimen geometry was Ø 10 mm x 15 mm and specimens were prepared from plane-parallel discs by wire erosion. The compression tests were carried out on the Gleeble 3800 GTC forming simulator. Under vacuum, the specimens were heated conductively at a heating rate of 4.5 K/s and held at the target temperature for 30 s to realise a homogeneous temperature distribution. Since partially liquid phases are already formed at this temperature, the specimens are compressed in a quartz sleeve. The quartz sleeve shields the specimen so that no particles can damage the test chamber. The punches are made of solid hard metal, which has a very high heat resistance. For the strain rates of 1 s⁻¹ and 10 s⁻¹ the tests were displacement controlled and for 0.1 s⁻¹ strain controlled. The specimens were upset to a plastic strain of 0.7. Each parameter combination from Table 2 was repeated five times. For the temperature range below 1,300 °C up to room temperature, flow curves were also recorded from the same material batch by cylinder compression tests [14].

The experimental investigations of the RFW process were carried out on the Genius Plus RS30 from KUKA. On the spindle side, a cylinder made of 41Cr4 with a Ø 40 mm and a length of 160 mm was inserted. A cylinder made of C22.8 with a Ø 40 mm and a length of 210 mm was used on the sledge side. The temperature measurement was carried out with a type S thermocouple. According to Fig. 1 (a), eight type K thermocouples (TC) were attached to the C22.8 bolts. In order to measure the temperature in the contact surface of the friction partners TC1 and TC5 were guided directly into the end face via small grooves. At 90° to this, measurements were taken axially around the circumference at a distance of 5 mm, 10 mm and 15 mm from the end face (TC2-TC4 and TC6-TC8). This allows the heat dissipation along the specimen to be determined. In addition, force, displacement as well as the rotational speed are recorded during the experiments. The applied time-rotational speed and force-time profiles for the 1 parameter set and 2 parameter set is shown in Fig. 1 (b). In the friction phase, heat is generated by the rotation and the applied axial force. By stopping the spindle, the upsetting phase is initiated and the bond is created. Fig. 1 (a) shows the process at the beginning of the friction phase on the left and after the end of the upsetting phase on the right. The RFW tests were repeated five times using this 1 parameter set with a rotational speed of 1,500 min⁻¹ and a upsetting force of 180 kN. In addition, a test with increased rotational speed of 2,000 min⁻¹ and increased upsetting force of 220 kN was carried out for later model validation in 2 parameter set. One specimen was shortened and separated along the
longitudinal axis by wire erosion and prepared metallographically and measured with the Keyence VR-3200 Profilometer. By etching with 5 % nitric acid (HNO3), the structural changes in the friction welding zone was analysed.

<figure><img src="image_1607.png" /><figcaption>Fig. 1 (a) RFW process at the beginning of the friction phase and at the end of the upsetting phase as well as the schematic positioning of the thermocouples; (b) rotational speed and upsetting force time curve for the 1 parameter set and 2 parameter set experiments.</figcaption></figure>

## Numerical Investigation

The RFW process was setup as a 2D model in the simulation software Simufact forming v16. For this purpose, two cylinders were created as workpieces according to the dimensions of the experiments and axially aligned, see Fig. 2. The modified material models were assigned to the two workpieces according to the experimental arrangement. Upper, lower and support dies were defined as rigid bodies. A spring force was applied between the support die and the upper die to apply the upsetting force. The course of the spring force was determined from the experiments and assigned to the spring. The positions of both the support die and the lower die are stationary in the room in order to feed the spring force into the welding zone. The upper die is movable in z-direction. Particles have been attached to the welding zone on the C22.8 workpiece, with the same spacing as the thermocouples in the experiments, compare Fig. 1 (a), which allow the temperature to be tracked through the process. The friction model according to Tresca with a friction factor of m = 0.3 was used to model the contact between the two workpieces. To analyse the geometry of the weld flash, a subroutine was used which records outer contour of the workpieces within a text file from the last increment. In this way, the influence of the model parameters on the weld flash formation can be precisely observed and compared. The friction welding module of Simufact forming allows the use of a 2D process, as the rotational movement and the resulting frictional heat are calculated via a substitute model. This significantly reduces the simulation time and simplifies the kinematic mapping of the process.
longitudinal axis by wire erosion and prepared metallographically and measured with the Keyence VR-3200 Profilometer. By etching with 5 % nitric acid (HNO3), the structural changes in the friction welding zone was analysed.

<figure><img src="image_1608.png" /><figcaption>Fig. 1 (a) RFW process at the beginning of the friction phase and at the end of the upsetting phase as well as the schematic positioning of the thermocouples; (b) rotational speed and upsetting force time curve for the 1 parameter set and 2 parameter set experiments.</figcaption></figure>

## Numerical Investigation

The RFW process was setup as a 2D model in the simulation software Simufact forming v16. For this purpose, two cylinders were created as workpieces according to the dimensions of the experiments and axially aligned, see Fig. 2. The modified material models were assigned to the two workpieces according to the experimental arrangement. Upper, lower and support dies were defined as rigid bodies. A spring force was applied between the support die and the upper die to apply the upsetting force. The course of the spring force was determined from the experiments and assigned to the spring. The positions of both the support die and the lower die are stationary in the room in order to feed the spring force into the welding zone. The upper die is movable in z-direction. Particles have been attached to the welding zone on the C22.8 workpiece, with the same spacing as the thermocouples in the experiments, compare Fig. 1 (a), which allow the temperature to be tracked through the process. The friction model according to Tresca with a friction factor of m = 0.3 was used to model the contact between the two workpieces. To analyse the geometry of the weld flash, a subroutine was used which records outer contour of the workpieces within a text file from the last increment. In this way, the influence of the model parameters on the weld flash formation can be precisely observed and compared. The friction welding module of Simufact forming allows the use of a 2D process, as the rotational movement and the resulting frictional heat are calculated via a substitute model. This significantly reduces the simulation time and simplifies the kinematic mapping of the process.
<figure><img src="image_1608.png" /><figcaption>Fig. 2 Configuration of the FE model and position of the refinement boxes.</figcaption></figure>

The frictional heat calculated by the equivalent model depends on the normal force, the rotational speed and the Y-factor. The Y-factor may range from 0.0 to 1.0 and indicates what percentage of the calculated frictional energy is introduced into the process [15]. The Y-factors can be specified depending on the temperature. In this way, for example, a change of frictional heat generation near the melting point can be taken into account. Due to the doughy/liquid state, no further frictional energy is generated in the real process until the temperature drops again below the melting point of approximately 1,490 °C. This can be realised by a Y-factor of 0.0 at the melting temperature of the material. Zhang et al. used Simufact forming for a friction stir rivet welding process and stated that they had used a temperature-dependent Y-factor for AA6061-T6 sheet [16]. Kubit and Trzepiecinski also use the friction welding module of Simufact forming to model friction stir spot welding Alclad 7075-T6 aluminium alloy sheets. They stated that they use a temperature-dependent Y-factor [17]. In this work, the Y-factor was also modeled as a function of temperature, see Table 3. The Y-factors were determined iteratively by comparing numerically determined and experimentally measured temperature curves. The resulting temperature curves were compared with the experimental curves of the 1 parameter set, allowing the appropriate Y-factors to be determined.

<table><caption>Table 3. Values of the examined Y-factors.</caption><tr><td>Y-factor</td><td>0.7</td><td>0.7</td><td>0.5</td><td>0.45</td><td>0.233</td><td>0.0</td></tr><tr><td>temperature</td><td>20 °C</td><td>300 °C</td><td>600 °C</td><td>1,200 °C</td><td>1,400 °C</td><td>1,490 °C</td></tr></table>

The workpieces are meshed with elements of the type Quads (10). To ensure that the results are independent of the selected element size, the element size was varied from 6 mm to 0.25 mm. The required calculation time increases exponentially with decreasing element size from 2 mm and reaches its maximum value of almost 350 min with an element size of 0.25 mm, see Fig. 3 (a).
<figure><img src="image_1609.png" /><figcaption>Fig. 2 Configuration of the FE model and position of the refinement boxes.</figcaption></figure>

The frictional heat calculated by the equivalent model depends on the normal force, the rotational speed and the Y-factor. The Y-factor may range from 0.0 to 1.0 and indicates what percentage of the calculated frictional energy is introduced into the process [15]. The Y-factors can be specified depending on the temperature. In this way, for example, a change of frictional heat generation near the melting point can be taken into account. Due to the doughy/liquid state, no further frictional energy is generated in the real process until the temperature drops again below the melting point of approximately 1,490 °C. This can be realised by a Y-factor of 0.0 at the melting temperature of the material. Zhang et al. used Simufact forming for a friction stir rivet welding process and stated that they had used a temperature-dependent Y-factor for AA6061-T6 sheet [16]. Kubit and Trzepiecinski also use the friction welding module of Simufact forming to model friction stir spot welding Alclad 7075-T6 aluminium alloy sheets. They stated that they use a temperature-dependent Y-factor [17]. In this work, the Y-factor was also modeled as a function of temperature, see Table 3. The Y-factors were determined iteratively by comparing numerically determined and experimentally measured temperature curves. The resulting temperature curves were compared with the experimental curves of the 1 parameter set, allowing the appropriate Y-factors to be determined.

<table><caption>Table 3. Values of the examined Y-factors.</caption><tr><td>Y-factor</td><td>0.7</td><td>0.7</td><td>0.5</td><td>0.45</td><td>0.233</td><td>0.0</td></tr><tr><td>temperature</td><td>20 °C</td><td>300 °C</td><td>600 °C</td><td>1,200 °C</td><td>1,400 °C</td><td>1,490 °C</td></tr></table>

The workpieces are meshed with elements of the type Quads (10). To ensure that the results are independent of the selected element size, the element size was varied from 6 mm to 0.25 mm. The required calculation time increases exponentially with decreasing element size from 2 mm and reaches its maximum value of almost 350 min with an element size of 0.25 mm, see Fig. 3 (a).
<figure><img src="image_1609.png" /><figcaption>Fig. 3 Results of the mesh study: (a) computation time over mesh size; (b) development of the maximum temperatures and the sledge path over the mesh size.</figcaption></figure>

The calculated maximum temperatures at the particles TC1-TC4 (squares) increase continuously until they stagnate at an element size of 1 mm and show convergence, Fig. 3 (b). The development of the maximum slide path (circle) over the element size is similar. Due to the comparatively long workpieces and the very localised deformation near the joining zone refinement boxes (RB, triangle) where used see Fig. 2. The workpiece was meshed with 2 mm element size and refined to 1 mm and 0.25 mm using two RB, coupled to the movements of the workpieces. By using the RB, the required calculation time can be reduced from 350 min to 35 min see Fig. 3 (a) and (b), thus saving resources.

## Results and Discussion

The recorded flow properties at 1,300 °C and 1,400 °C are shown in Fig. 4. Both materials show similar flow characteristics. The flow stress decreases with increasing temperature, whereas the flow stress increases with increasing strain rate. In the temperature range investigated, the flow stress for the C22.8 is slightly higher than the flow stress of the 41Cr4. With the determined flow curves, the existing material models were extended so that the flow behaviour for C22.8 and 41Cr4 from 20 °C to 1,400 °C can now be de-scribed and used in the FE simulation.

<figure><img src="image_1610.png" /><figcaption>Fig. 4. Flow curves for (a) C22.8 and (b) 41Cr4 at 1,300 °C and 1,400 °C.</figcaption></figure>

The Y-factor-temperature curve was fitted in such a way that the best possible agreement could be achieved with the experimentally recorded temperature curves, sledge path and weld flash geometry of the 1 parameter set. In Fig. 5 (a)-(d) the experimental temperature curves of the
<figure><img src="image_1610.png" /><figcaption>Fig. 3 Results of the mesh study: (a) computation time over mesh size; (b) development of the maximum temperatures and the sledge path over the mesh size.</figcaption></figure>

The calculated maximum temperatures at the particles TC1-TC4 (squares) increase continuously until they stagnate at an element size of 1 mm and show convergence, Fig. 3 (b). The development of the maximum slide path (circle) over the element size is similar. Due to the comparatively long workpieces and the very localised deformation near the joining zone refinement boxes (RB, triangle) where used see Fig. 2. The workpiece was meshed with 2 mm element size and refined to 1 mm and 0.25 mm using two RB, coupled to the movements of the workpieces. By using the RB, the required calculation time can be reduced from 350 min to 35 min see Fig. 3 (a) and (b), thus saving resources.

## Results and Discussion

The recorded flow properties at 1,300 °C and 1,400 °C are shown in Fig. 4. Both materials show similar flow characteristics. The flow stress decreases with increasing temperature, whereas the flow stress increases with increasing strain rate. In the temperature range investigated, the flow stress for the C22.8 is slightly higher than the flow stress of the 41Cr4. With the determined flow curves, the existing material models were extended so that the flow behaviour for C22.8 and 41Cr4 from 20 °C to 1,400 °C can now be de-scribed and used in the FE simulation.

<figure><img src="image_1611.png" /><figcaption>Fig. 4. Flow curves for (a) C22.8 and (b) 41Cr4 at 1,300 °C and 1,400 °C.</figcaption></figure>

The Y-factor-temperature curve was fitted in such a way that the best possible agreement could be achieved with the experimentally recorded temperature curves, sledge path and weld flash geometry of the 1 parameter set. In Fig. 5 (a)-(d) the experimental temperature curves of the
respective thermocouples are compared to the simulated curves. The temperature curves of the axially attached particles lie within the scatter range of the experimental curves (Fig. 6 (b)-(d)). For TC1 and TC5, the experimental curves scatter very strongly. The reason for this is the high mechanical load in the front surface during the RFW process, which influences the results of the temperature measurement. Therefore, the temperature for TC1/TC5 is higher in the simulation, but remains below the melting temperature of 1,492 °C. A comparison of the simulated and experimental sledge path shows no deviations during the friction phase. The duration of the upsetting phase as well as the steep increase of the sledge path is also well predicted with the FE model Fig. 6 (e). However, the maximum sledge path is underestimated by approx. 2 mm in the simulation. In addition, the simulated sledge path runs constant after reaching the maximum, while in the experiment the sledge path continues to increase. Possible reasons for the deviations could be due to the machine stiffness or to slipping of the specimen in the clamping jaws of the sledge. These effects are not taken into account in the simulation. In Fig. 6 (f), a friction welded specimen is shown in longitudinal section, with the simulated contours superimposed. The specimen was etched so that the joining zone and the heat-affected zone are visible. With the simulated flash geometries, the flash geometries of the experiments were well predicted. Only in the case of 41Cr4 slight deviations in the radii and in the transition from the initial diameter to the weld flash are visible. Apart from the deviation at TC1 and the maximum sledge path, a successful model verification was achieved, which predicts the experimental results well.

<figure><img src="image_1611.png" /><figcaption>Fig. 5 Comparison of the 1 parameter set of the experimental and simulated results [18] (a) temperature-time curves of TC1 and TC5; (b) temperature-time curves of TC2 and TC6; (c) temperature-time curves of TC3 and TC7; (d) temperature-time curves of TC4 and TC8; (e) Sledge path time curves; (f) weld flash geometry of C22.8 and 41Cr4.</figcaption></figure>

Finally, a simulation with the 2 parameter set was carried out with the verified Y-factor-temperature curve in order to examine the general validity of the determined Y-temperature curve. During the 2 parameter set experiment, the rotational speed in the friction phase was increased to 2,000 min⁻¹ and the upsetting force in the upsetting phase was increased to 220 kN, see Fig. 1 (b). Since the substitute model is not only based on the Y-factor but also on the rotational speed and the normal stress, it can be proofed whether the determined Y-factors also provide sufficiently
respective thermocouples are compared to the simulated curves. The temperature curves of the axially attached particles lie within the scatter range of the experimental curves (Fig. 6 (b)-(d)). For TC1 and TC5, the experimental curves scatter very strongly. The reason for this is the high mechanical load in the front surface during the RFW process, which influences the results of the temperature measurement. Therefore, the temperature for TC1/TC5 is higher in the simulation, but remains below the melting temperature of 1,492 °C. A comparison of the simulated and experimental sledge path shows no deviations during the friction phase. The duration of the upsetting phase as well as the steep increase of the sledge path is also well predicted with the FE model Fig. 6 (e). However, the maximum sledge path is underestimated by approx. 2 mm in the simulation. In addition, the simulated sledge path runs constant after reaching the maximum, while in the experiment the sledge path continues to increase. Possible reasons for the deviations could be due to the machine stiffness or to slipping of the specimen in the clamping jaws of the sledge. These effects are not taken into account in the simulation. In Fig. 6 (f), a friction welded specimen is shown in longitudinal section, with the simulated contours superimposed. The specimen was etched so that the joining zone and the heat-affected zone are visible. With the simulated flash geometries, the flash geometries of the experiments were well predicted. Only in the case of 41Cr4 slight deviations in the radii and in the transition from the initial diameter to the weld flash are visible. Apart from the deviation at TC1 and the maximum sledge path, a successful model verification was achieved, which predicts the experimental results well.

<figure><img src="image_1612.png" /><figcaption>Fig. 5 Comparison of the 1 parameter set of the experimental and simulated results [18] (a) temperature-time curves of TC1 and TC5; (b) temperature-time curves of TC2 and TC6; (c) temperature-time curves of TC3 and TC7; (d) temperature-time curves of TC4 and TC8; (e) Sledge path time curves; (f) weld flash geometry of C22.8 and 41Cr4.</figcaption></figure>

Finally, a simulation with the 2 parameter set was carried out with the verified Y-factor-temperature curve in order to examine the general validity of the determined Y-temperature curve. During the 2 parameter set experiment, the rotational speed in the friction phase was increased to 2,000 min⁻¹ and the upsetting force in the upsetting phase was increased to 220 kN, see Fig. 1 (b). Since the substitute model is not only based on the Y-factor but also on the rotational speed and the normal stress, it can be proofed whether the determined Y-factors also provide sufficiently
accurate results for varying process parameters. In Fig. 7 (a)-(d) the simulated and experimental temperature curves are compared. At all measuring points the simulation overestimates the temperature curves from the experiment. Especially in the friction phase, the temperature at TC1 rises steeply, resulting in a high amount of frictional heat in the contact zone, which is conducted into the workpiece. In the upsetting phase, on the other hand, cooling is well achieved. The increased temperature during the upsetting phase is due to the increased rotational speed, since the increase in upsetting force only takes effect in the upsetting phase. Due to the fact that too much frictional heat is calculated by the substitute model, the material already begins to flow during the friction phase, which leads to an overestimation of the sledge path in the friction phase, see Fig. 7 (e). As a result, the maximum sledge path of the simulation and the experiment are equal in the upsetting phase. Although the maximum sledge paths are identical, the simulated weld flash geometries clearly overestimate the real flash formation for both materials, see Fig. 7 (f). Due to the effects of machine stiffness or slipping of the workpiece, which are not shown, the maximum sledge path should not be used as the basis for evaluation, but rather the sledge path at the end of the friction phase. A successful validation of the determined Y-factors on changed process parameters is therefore not given. For this purpose, further investigations must be carried out on the influence of the rotational speed and the upsetting force in combination with the Y-factor during the friction phase.

<figure><img src="image_1612.png" /><figcaption>Fig. 6 Comparison of the 2 parameter set of the experimental and simulated results [19] (a) temperature-time curves of TC1 and TC5; (b) temperature-time curves of TC2 and TC6; (c) temperature-time curves of TC3 and TC7; (d) temperature-time curves of TC4 and TC8; (e) sledge path time curves; (f) weld flash geometry of C22.8 and 41Cr4.</figcaption></figure>

## Summary and Conclusions

In this article, the RFW process was modelled by means of FEM using Simufact forming. The focus was on the calculation of the frictional heat via a substitute model which is dependent on the Y-factor, the rotational speed and the normal stress. First of all, the existing material models for C22.8 and 41Cr4 were extended for a temperature range up to 1,400 °C by means of cylinder compression tests. After the mesh study was carried out, the influence of the Y-factor on the temperature, the sledge path as well as the weld flash geometry was investigated. A Y-temperature curve was fitted to the results of the experiments of the 1 parameter set and verified. Finally, the
accurate results for varying process parameters. In Fig. 7 (a)-(d) the simulated and experimental temperature curves are compared. At all measuring points the simulation overestimates the temperature curves from the experiment. Especially in the friction phase, the temperature at TC1 rises steeply, resulting in a high amount of frictional heat in the contact zone, which is conducted into the workpiece. In the upsetting phase, on the other hand, cooling is well achieved. The increased temperature during the upsetting phase is due to the increased rotational speed, since the increase in upsetting force only takes effect in the upsetting phase. Due to the fact that too much frictional heat is calculated by the substitute model, the material already begins to flow during the friction phase, which leads to an overestimation of the sledge path in the friction phase, see Fig. 7 (e). As a result, the maximum sledge path of the simulation and the experiment are equal in the upsetting phase. Although the maximum sledge paths are identical, the simulated weld flash geometries clearly overestimate the real flash formation for both materials, see Fig. 7 (f). Due to the effects of machine stiffness or slipping of the workpiece, which are not shown, the maximum sledge path should not be used as the basis for evaluation, but rather the sledge path at the end of the friction phase. A successful validation of the determined Y-factors on changed process parameters is therefore not given. For this purpose, further investigations must be carried out on the influence of the rotational speed and the upsetting force in combination with the Y-factor during the friction phase.

<figure><img src="image_1613.png" /><figcaption>Fig. 6 Comparison of the 2 parameter set of the experimental and simulated results [19] (a) temperature-time curves of TC1 and TC5; (b) temperature-time curves of TC2 and TC6; (c) temperature-time curves of TC3 and TC7; (d) temperature-time curves of TC4 and TC8; (e) sledge path time curves; (f) weld flash geometry of C22.8 and 41Cr4.</figcaption></figure>

## Summary and Conclusions

In this article, the RFW process was modelled by means of FEM using Simufact forming. The focus was on the calculation of the frictional heat via a substitute model which is dependent on the Y-factor, the rotational speed and the normal stress. First of all, the existing material models for C22.8 and 41Cr4 were extended for a temperature range up to 1,400 °C by means of cylinder compression tests. After the mesh study was carried out, the influence of the Y-factor on the temperature, the sledge path as well as the weld flash geometry was investigated. A Y-temperature curve was fitted to the results of the experiments of the 1 parameter set and verified. Finally, the
Y-temperature curve was tested regarding the validity for the process parameters of the 2 parameter set. From this work it can be concluded that the RFW can be reproduced with the substitute model in Simufact forming. By fitting a Y-factor-temperature curve to the experimentally recorded data, it was possible to correctly predict the RFW numerically. However, this Y-factor-temperature curve is not generally valid, as the comparison with further experimental data has shown. Thus, it is possible to numerically predict the RFW correctly if the experimental data is available, but the model is currently not suitable for carrying out variations of process parameters.

Since the mathematical structure of the substitute model is not known, the exact mathematical relationships are unknown. Therefore, the frictional heat calculation should be determined by subroutines depending on the friction present and not via an substitute model where the exact physical relationships are unknown.

## Acknowledgments

The results presented in this paper were obtained within the Collaborative Research Center 1153 "Process chain to produce hybrid high performance components by Tailored Forming" in the subproject A01 and B03, funded by the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation)—252662854. The authors thank the German Research Foundation (DFG) for their financial support of this project.

## References

[1] U. Dilthey, Schweißtechnische Fertigungsverfahren 2: Verhalten der Werkstoffe beim Schweißen; Springer-Verlag: Berlin/Heidelberg, 2005, ISBN 3-540-21674-X.

[2] T. Lienert, T. Siewert, S. Babu, V. Acoff, ASM handbook - Welding Fundamentals and Processes, 2nd print; ASM International: Metals Park, Ohio, 2015, ISBN 978-1-61503-133-7.

[3] W. Li, A. Vairis, M. Preuss, T. Ma, Linear and rotary friction welding review. International Materials Reviews 2016, 61, pp. 71–100. https://doi.org/10.1080/09506608.2015.1109214

[4] M. Soucail, A. Moal, L. Naze, E. Massoni, C. Levaillant, Y. Bienvenu, Microstructural Study and Numerical Simulation of Inertia Friction Welding of Astroloy. In Superalloys 1992 (Seventh International Symposium). Superalloys, 20–24 Sep. 1992; TMS, 1992 - 1992; pp. 847–856, ISBN 0-87339-189-6

[5] T.C. Nguyen, D.C. Weckman, A thermal and microstructure evolution model of direct-drive friction welding of plain carbon steel. Metall and Materi Trans B 2006, 37, pp. 275–292. https://doi.org/10.1007/BF02693157

[6] D. Schmicker, A holistic approach on the simulation of rotary friction welding; Berlin, 2015, ISBN 3737575177.

[7] H.T. My Nu, T.T. Le, L.P. Minh, N.H. Loc, A Study on Rotary Friction Welding of Titanium Alloy (Ti6Al4V). Advances in Materials Science and Engineering 2019, 2019, pp. 1–9. https://doi.org/10.1155/2019/4728213

[8] F.F. Wang, W.Y. Li, J.L. Li, A. Vairis, Process parameter analysis of inertia friction welding nickel-based superalloy. Int J Adv Manuf Technol 2014, 71, pp. 1909–1918. https://doi.org/10.1007/s00170-013-5569-6

[9] P. Li, J. Li, H. Dong, Analytical description of heat generation and temperature field during the initial stage of rotary friction welding. Journal of Manufacturing Processes 2017, 25, pp. 181–184. https://doi.org/10.1016/j.jmapro.2016.12.003
Y-temperature curve was tested regarding the validity for the process parameters of the 2 parameter set. From this work it can be concluded that the RFW can be reproduced with the substitute model in Simufact forming. By fitting a Y-factor-temperature curve to the experimentally recorded data, it was possible to correctly predict the RFW numerically. However, this Y-factor-temperature curve is not generally valid, as the comparison with further experimental data has shown. Thus, it is possible to numerically predict the RFW correctly if the experimental data is available, but the model is currently not suitable for carrying out variations of process parameters.

Since the mathematical structure of the substitute model is not known, the exact mathematical relationships are unknown. Therefore, the frictional heat calculation should be determined by subroutines depending on the friction present and not via an substitute model where the exact physical relationships are unknown.

## Acknowledgments

The results presented in this paper were obtained within the Collaborative Research Center 1153 "Process chain to produce hybrid high performance components by Tailored Forming" in the subproject A01 and B03, funded by the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation)—252662854. The authors thank the German Research Foundation (DFG) for their financial support of this project.

## References

[1] U. Dilthey, Schweißtechnische Fertigungsverfahren 2: Verhalten der Werkstoffe beim Schweißen; Springer-Verlag: Berlin/Heidelberg, 2005, ISBN 3-540-21674-X.

[2] T. Lienert, T. Siewert, S. Babu, V. Acoff, ASM handbook - Welding Fundamentals and Processes, 2nd print; ASM International: Metals Park, Ohio, 2015, ISBN 978-1-61503-133-7.

[3] W. Li, A. Vairis, M. Preuss, T. Ma, Linear and rotary friction welding review. International Materials Reviews 2016, 61, pp. 71–100. https://doi.org/10.1080/09506608.2015.1109214

[4] M. Soucail, A. Moal, L. Naze, E. Massoni, C. Levaillant, Y. Bienvenu, Microstructural Study and Numerical Simulation of Inertia Friction Welding of Astroloy. In Superalloys 1992 (Seventh International Symposium). Superalloys, 20–24 Sep. 1992; TMS, 1992 - 1992; pp. 847–856, ISBN 0-87339-189-6

[5] T.C. Nguyen, D.C. Weckman, A thermal and microstructure evolution model of direct-drive friction welding of plain carbon steel. Metall and Materi Trans B 2006, 37, pp. 275–292. https://doi.org/10.1007/BF02693157

[6] D. Schmicker, A holistic approach on the simulation of rotary friction welding; Berlin, 2015, ISBN 3737575177.

[7] H.T. My Nu, T.T. Le, L.P. Minh, N.H. Loc, A Study on Rotary Friction Welding of Titanium Alloy (Ti6Al4V). Advances in Materials Science and Engineering 2019, 2019, pp. 1–9. https://doi.org/10.1155/2019/4728213

[8] F.F. Wang, W.Y. Li, J.L. Li, A. Vairis, Process parameter analysis of inertia friction welding nickel-based superalloy. Int J Adv Manuf Technol 2014, 71, pp. 1909–1918. https://doi.org/10.1007/s00170-013-5569-6

[9] P. Li, J. Li, H. Dong, Analytical description of heat generation and temperature field during the initial stage of rotary friction welding. Journal of Manufacturing Processes 2017, 25, pp. 181–184. https://doi.org/10.1016/j.jmapro.2016.12.003
[10] X. Nan, J. Xiong, F. Jin, X. Li, Z. Liao, F. Zhang, J. Li, Modeling of rotary friction welding process based on maximum entropy production principle. Journal of Manufacturing Processes 2019, 37, pp. 21-27. https://doi.org/10.1016/j.jmapro.2018.11.016

[11] B.-A. Behrens, J. Uhe, Introduction to tailored forming. Prod. Eng. Res. Devel. 2021, 15, pp. 133-136. https://doi.org/10.1007/s11740-021-01022-w

[12] B.-A. Behrens, J. Uhe, T. Petersen, F. Nürnberger, C. Kahra, I. Ross, R. Laeger, Contact Geometry Modification of Friction-Welded Semi-Finished Products to Improve the Bonding of Hybrid Components. Metals 2021, 11. https://doi.org/10.3390/met11010115

[13] B.-A. Behrens, D. Duran, T. Matthias, I. Ross, Enhancement of the interface of friction welded steel-aluminium joints. Prod. Eng. Res. Devel. 2021, 15, pp. 169-176. https://doi.org/10.1007/s11740-020-00994-5

[14] B.-A. Behrens, A. Bouguecha, C. Bonk, T. Matthias, Importance of material and friction characterisation for FE-aided process design of hybrid bevel gears. In. Proceedings of the International Conference of Global Network for Innovative Technology and AWAM International Conference in Civil Engineering (IGNITE-AICCE'17): Sustainable Technology and Practice for Infrastructure and Community Resilience, Penang, Malaysia, 8-9 August 2017; pp. 190016.

[15] N. Schubert, A. Sterzig, R. Mauermann, S. Hilbers, P. Kolbe, C. Kuhn, Development of a Friction Welded Joint for Future Industrial Application. MSF 2019, 949, pp. 119-124. https://doi.org/10.4028/www.scientific.net/MSF.949.119

[16] P. Zhang, S. Zhao, W. Wang, H. Zhang, J. Zhang, C. Yang, Y. Wang, W. Wei, G. Ma, Numerical Simulation and Experimental Investigation of Friction Stir Rivet Welding Process for AA6061-T6. dtmse 2021. https://doi.org/10.12783/dtmse/ameme2020/35546

[17] A. Kubit, T. Trzepiecinski, A fully coupled thermo-mechanical numerical modelling of the refill friction stir spot welding process in Alclad 7075-T6 aluminium alloy sheets. Archiv.Civ.Mech.Eng 2020, 2020. https://doi.org/10.1007/s43452-020-00127-w

[18] N. Mohnfeld, Dataset: Temperature measurement Rotary friction welding 1setup. 2023, https://doi.org/10.25835/r2534sqw

[19] N. Mohnfeld, Dataset: Temperature measurement Rotary friction welding 2setup. 2024, https://doi.org/10.25835/2jc9g610
[10] X. Nan, J. Xiong, F. Jin, X. Li, Z. Liao, F. Zhang, J. Li, Modeling of rotary friction welding process based on maximum entropy production principle. Journal of Manufacturing Processes 2019, 37, pp. 21-27. https://doi.org/10.1016/j.jmapro.2018.11.016

[11] B.-A. Behrens, J. Uhe, Introduction to tailored forming. Prod. Eng. Res. Devel. 2021, 15, pp. 133-136. https://doi.org/10.1007/s11740-021-01022-w

[12] B.-A. Behrens, J. Uhe, T. Petersen, F. Nürnberger, C. Kahra, I. Ross, R. Laeger, Contact Geometry Modification of Friction-Welded Semi-Finished Products to Improve the Bonding of Hybrid Components. Metals 2021, 11. https://doi.org/10.3390/met11010115

[13] B.-A. Behrens, D. Duran, T. Matthias, I. Ross, Enhancement of the interface of friction welded steel-aluminium joints. Prod. Eng. Res. Devel. 2021, 15, pp. 169-176. https://doi.org/10.1007/s11740-020-00994-5

[14] B.-A. Behrens, A. Bouguecha, C. Bonk, T. Matthias, Importance of material and friction characterisation for FE-aided process design of hybrid bevel gears. In. Proceedings of the International Conference of Global Network for Innovative Technology and AWAM International Conference in Civil Engineering (IGNITE-AICCE'17): Sustainable Technology and Practice for Infrastructure and Community Resilience, Penang, Malaysia, 8-9 August 2017; pp. 190016.

[15] N. Schubert, A. Sterzig, R. Mauermann, S. Hilbers, P. Kolbe, C. Kuhn, Development of a Friction Welded Joint for Future Industrial Application. MSF 2019, 949, pp. 119-124. https://doi.org/10.4028/www.scientific.net/MSF.949.119

[16] P. Zhang, S. Zhao, W. Wang, H. Zhang, J. Zhang, C. Yang, Y. Wang, W. Wei, G. Ma, Numerical Simulation and Experimental Investigation of Friction Stir Rivet Welding Process for AA6061-T6. dtmse 2021. https://doi.org/10.12783/dtmse/ameme2020/35546

[17] A. Kubit, T. Trzepiecinski, A fully coupled thermo-mechanical numerical modelling of the refill friction stir spot welding process in Alclad 7075-T6 aluminium alloy sheets. Archiv.Civ.Mech.Eng 2020, 2020. https://doi.org/10.1007/s43452-020-00127-w

[18] N. Mohnfeld, Dataset: Temperature measurement Rotary friction welding 1setup. 2023, https://doi.org/10.25835/r2534sqw

[19] N. Mohnfeld, Dataset: Temperature measurement Rotary friction welding 2setup. 2024, https://doi.org/10.25835/2jc9g610
# Influence of cold rolled surface structures with undercuts for interlocking joints on bending processes

RINGEL Aron¹,ᵃ* and BAILLY David¹,ᵇ

¹Institute of Metal Forming (IBF), RWTH Aachen University, Intzestraße 10, 52072 Aachen, Germany

ᵃaron.ringel@ibf.rwth-aachen.de, ᵇdavid.bailly@ibf.rwth-aachen.de

**Keywords:** Interlocking, Multi-Material, Structured Cold Rolling, 3-Point Bending, FE-Modelling

**Abstract.** Lightweight design is one of the methods to reduce CO₂ emissions and optimize energy efficiency in the transportation sector. The main motivation of this study arise from weight reduction through multi-material design with the materials commonly used in the automotive industry, steel and aluminum. Metallurgical bonding of steel and aluminum carries the risk of forming brittle intermetallic phases. Hence, new joining techniques such as joining by forming or casting are promising for these multi-material components. Previously, a hybrid component method was presented using channel-like surface structures with undercuts on a steel sheet created in a modified cold rolling process. In a subsequent high-pressure die casting, the channels were filled with aluminum melt, forming an interlocking connection as it solidified. As automotive components demand increased complexity, the bending of the structured sheet before casting was investigated. This study aims to analyze how the surface structure affects the bending process. Numerical simulations and experiments were used to investigate the effect on the maximum bending force, the resulting bending angle and springback. Therefore, the parameters bending angle, bending radius, and the lateral or longitudinal orientations of the channel structure on either sides of the bend were taken into account. The results showed a strong influence of the lateral and longitudinal orientation on the maximum bending force. Furthermore, a minor effect of the bending radius on the force and springback was found.

**Introduction**

To achieve the needed reduction of CO₂ emissions and improve the energy efficiency, changes in the transportation sector are needed. Lightweight design is an established method to optimize the energy efficiency, which has still a huge potential, especially for the automotive sector and in the targeted transformation towards electric vehicles. One way to reduce weight is by exploiting the advantages of different materials. Therefore, this research focuses on a multi-material design of two commonly used materials in the automotive industry. A steel is used for its high strength and ductility in combination with aluminum, which provides high specific strength and exceptional lightweight potential. A component of this combination can meet the automotive key requirements, i.e. performance, function and safety, while reducing weight.

**State of the Art**

The joining of different metal classes, in this study steel and aluminum, is challenging due to their dissimilar thermo-physical properties [1]. A metallurgical bonding brings the risk of the formation of brittle intermetallic phases and makes joining techniques such as fusion welding unsuitable. Therefore, multi-material components require new joining methods like joining by forming or casting [2,3]. Advanced processes like cold or friction stir welding can form a metallurgical bond under high interfacial pressure overcoming the problems of fusion welding. Mechanical joining processes like riveting or clinching interlock different materials by forming and avoid a

<figure><img src="image_1613.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.
# Influence of cold rolled surface structures with undercuts for interlocking joints on bending processes

RINGEL Aron¹,ᵃ* and BAILLY David¹,ᵇ

¹Institute of Metal Forming (IBF), RWTH Aachen University, Intzestraße 10, 52072 Aachen, Germany

ᵃaron.ringel@ibf.rwth-aachen.de, ᵇdavid.bailly@ibf.rwth-aachen.de

**Keywords:** Interlocking, Multi-Material, Structured Cold Rolling, 3-Point Bending, FE-Modelling

**Abstract.** Lightweight design is one of the methods to reduce CO₂ emissions and optimize energy efficiency in the transportation sector. The main motivation of this study arise from weight reduction through multi-material design with the materials commonly used in the automotive industry, steel and aluminum. Metallurgical bonding of steel and aluminum carries the risk of forming brittle intermetallic phases. Hence, new joining techniques such as joining by forming or casting are promising for these multi-material components. Previously, a hybrid component method was presented using channel-like surface structures with undercuts on a steel sheet created in a modified cold rolling process. In a subsequent high-pressure die casting, the channels were filled with aluminum melt, forming an interlocking connection as it solidified. As automotive components demand increased complexity, the bending of the structured sheet before casting was investigated. This study aims to analyze how the surface structure affects the bending process. Numerical simulations and experiments were used to investigate the effect on the maximum bending force, the resulting bending angle and springback. Therefore, the parameters bending angle, bending radius, and the lateral or longitudinal orientations of the channel structure on either sides of the bend were taken into account. The results showed a strong influence of the lateral and longitudinal orientation on the maximum bending force. Furthermore, a minor effect of the bending radius on the force and springback was found.

**Introduction**

To achieve the needed reduction of CO₂ emissions and improve the energy efficiency, changes in the transportation sector are needed. Lightweight design is an established method to optimize the energy efficiency, which has still a huge potential, especially for the automotive sector and in the targeted transformation towards electric vehicles. One way to reduce weight is by exploiting the advantages of different materials. Therefore, this research focuses on a multi-material design of two commonly used materials in the automotive industry. A steel is used for its high strength and ductility in combination with aluminum, which provides high specific strength and exceptional lightweight potential. A component of this combination can meet the automotive key requirements, i.e. performance, function and safety, while reducing weight.

**State of the Art**

The joining of different metal classes, in this study steel and aluminum, is challenging due to their dissimilar thermo-physical properties [1]. A metallurgical bonding brings the risk of the formation of brittle intermetallic phases and makes joining techniques such as fusion welding unsuitable. Therefore, multi-material components require new joining methods like joining by forming or casting [2,3]. Advanced processes like cold or friction stir welding can form a metallurgical bond under high interfacial pressure overcoming the problems of fusion welding. Mechanical joining processes like riveting or clinching interlock different materials by forming and avoid a

<figure><img src="image_1614.png" /></figure>

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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_1614.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