gauge length lines return to be parallel throughout the test and the fracture occurs when the molecular chains reach a certain length.

<figure><img src="image_1532.png" /><figcaption>Fig. 2. (a) Stress triaxiality distribution over the thickness of a notched sample (R5 mm) under elastic and elastic plastic conditions and (b) mechanical behaviour of smooth sample.</figcaption></figure>

PEEK is temperature sensitive, and raising the temperature below the glass transition temperature enhances its formability. To investigate the effect of temperature on stress triaxiality, tensile tests on smooth and notched samples were performed at various temperatures, including room temperature (22 °C), 80 °C, 120 °C, and 150 °C. The influence of temperature on the stress triaxiality of a notched sample (R5mm) is shown in Fig. 3 (a). It can be noted that the stress triaxiality decreases as temperature rises, which indicates that as PEEK formability increases with temperature, the stress triaxiality decreases. Temperature has the same effect on stress triaxiality in all notched samples, and the effect of temperature on smooth samples is identical to the effect of plastic deformation; the stress triaxiality of smooth samples does not change with temperature. Table 2 shows the experimental fracture strain and stress triaxiality results at various temperatures. To clarify why stress triaxiality decreases with rising temperature in notched samples, Fig. 3 (b) compares the FE results of stresses distribution at necking area of R5 mm notched sample at room temperature with one at 150 °C. It is assumed that σ₁ is the tensile stress, which obtained from the experimental tensile test and decreases as the testing temperature rises. σ₂ and σ₃ denote the stress in the width and thickness directions, respectively. σ₂ and σ₃ are close to zero before necking and increase when necking begins. Because the yield stress of PEEK is greater at room temperature than at 150 °C, high values of σ₂ and σ₃ are required to compress the sample and induce necking in the sample at room temperature, whereas the required stresses (σ₂ and σ₃) are decreased at 150 °C as shown in the figure. When σ₂ and σ₃ are reduced, the hydrostatic stress decreases, resulting in a decrease in stress triaxiality.
gauge length lines return to be parallel throughout the test and the fracture occurs when the molecular chains reach a certain length.

<figure><img src="image_1536.png" /><figcaption>Fig. 2. (a) Stress triaxiality distribution over the thickness of a notched sample (R5 mm) under elastic and elastic plastic conditions and (b) mechanical behaviour of smooth sample.</figcaption></figure>

PEEK is temperature sensitive, and raising the temperature below the glass transition temperature enhances its formability. To investigate the effect of temperature on stress triaxiality, tensile tests on smooth and notched samples were performed at various temperatures, including room temperature (22 °C), 80 °C, 120 °C, and 150 °C. The influence of temperature on the stress triaxiality of a notched sample (R5mm) is shown in Fig. 3 (a). It can be noted that the stress triaxiality decreases as temperature rises, which indicates that as PEEK formability increases with temperature, the stress triaxiality decreases. Temperature has the same effect on stress triaxiality in all notched samples, and the effect of temperature on smooth samples is identical to the effect of plastic deformation; the stress triaxiality of smooth samples does not change with temperature. Table 2 shows the experimental fracture strain and stress triaxiality results at various temperatures. To clarify why stress triaxiality decreases with rising temperature in notched samples, Fig. 3 (b) compares the FE results of stresses distribution at necking area of R5 mm notched sample at room temperature with one at 150 °C. It is assumed that σ₁ is the tensile stress, which obtained from the experimental tensile test and decreases as the testing temperature rises. σ₂ and σ₃ denote the stress in the width and thickness directions, respectively. σ₂ and σ₃ are close to zero before necking and increase when necking begins. Because the yield stress of PEEK is greater at room temperature than at 150 °C, high values of σ₂ and σ₃ are required to compress the sample and induce necking in the sample at room temperature, whereas the required stresses (σ₂ and σ₃) are decreased at 150 °C as shown in the figure. When σ₂ and σ₃ are reduced, the hydrostatic stress decreases, resulting in a decrease in stress triaxiality.
<figure><img src="image_1533.png" /><figcaption>Fig. 3. Sensitivity of stress triaxiality (a) and stresses distribution (b) to temperature changes of R5 mm notched sample.</figcaption></figure>

<table><caption>Table 2. Fracture strain-stress triaxiality (η) of PEEK at different notches temperatures.</caption><thead><tr><th rowspan="2">Temp.<br/>[°C]</th><th colspan="2">Smooth</th><th colspan="2">R5</th><th colspan="2">R3</th><th colspan="2">R1</th><th colspan="2">R0.5</th></tr><tr><th>Strain</th><th>η</th><th>Strain</th><th>η</th><th>Strain</th><th>η</th><th>Strain</th><th>η</th><th>Strain</th><th>η</th></tr></thead><tbody><tr><td>22</td><td>1.29</td><td>0.33</td><td>0.52</td><td>0.93</td><td>0.38</td><td>1.04</td><td>0.15</td><td>1.20</td><td>0.07</td><td>1.37</td></tr><tr><td>80</td><td>1.55</td><td>0.33</td><td>0.62</td><td>0.79</td><td>0.59</td><td>0.96</td><td>0.16</td><td>1.04</td><td>0.09</td><td>1.28</td></tr><tr><td>120</td><td>1.86</td><td>0.33</td><td>0.78</td><td>0.73</td><td>0.67</td><td>0.78</td><td>0.22</td><td>0.93</td><td>0.12</td><td>1.19</td></tr><tr><td>150</td><td>1.94</td><td>0.33</td><td>1.33</td><td>0.55</td><td>0.96</td><td>0.60</td><td>0.35</td><td>0.86</td><td>0.24</td><td>1.03</td></tr></tbody></table>

## Determination of Hooputra ductile damage parameters

The Hooputra ductile damage model is utilised in this work to predict the fracture initiation and propagation in uniaxial tensile testing and the SPIF process. To calculate the Hooputra parameters at different temperatures the values of fracture strain and corresponding stress triaxiality from Table 2 are plotted as five isolated points at each temperature. To obtain the best match between Eq.1 and the five isolated points, a curve fitting procedure is carried out using MATLAB tools. The best-fitting values of a, b, and c represent the parameters of the Hooputra ductile damage model. Fig. 4 depicts the Hooputra damage curves at various temperatures based on data from Table 2, and Table 3 presents the relevant Hooputra parameters. According to Fig. 4 and Table 3, the temperature affects the Hooputra damage curve and Hooputra parameters; hence, the Hooputra ductile damage model should be modified to account for temperature impacts.
<figure><img src="image_1537.png" /><figcaption>Fig. 3. Sensitivity of stress triaxiality (a) and stresses distribution (b) to temperature changes of R5 mm notched sample.</figcaption></figure>

<table><caption>Table 2. Fracture strain-stress triaxiality (η) of PEEK at different notches temperatures.</caption><thead><tr><th rowspan="2">Temp.<br/>[°C]</th><th colspan="2">Smooth</th><th colspan="2">R5</th><th colspan="2">R3</th><th colspan="2">R1</th><th colspan="2">R0.5</th></tr><tr><th>Strain</th><th>η</th><th>Strain</th><th>η</th><th>Strain</th><th>η</th><th>Strain</th><th>η</th><th>Strain</th><th>η</th></tr></thead><tbody><tr><td>22</td><td>1.29</td><td>0.33</td><td>0.52</td><td>0.93</td><td>0.38</td><td>1.04</td><td>0.15</td><td>1.20</td><td>0.07</td><td>1.37</td></tr><tr><td>80</td><td>1.55</td><td>0.33</td><td>0.62</td><td>0.79</td><td>0.59</td><td>0.96</td><td>0.16</td><td>1.04</td><td>0.09</td><td>1.28</td></tr><tr><td>120</td><td>1.86</td><td>0.33</td><td>0.78</td><td>0.73</td><td>0.67</td><td>0.78</td><td>0.22</td><td>0.93</td><td>0.12</td><td>1.19</td></tr><tr><td>150</td><td>1.94</td><td>0.33</td><td>1.33</td><td>0.55</td><td>0.96</td><td>0.60</td><td>0.35</td><td>0.86</td><td>0.24</td><td>1.03</td></tr></tbody></table>

## Determination of Hooputra ductile damage parameters

The Hooputra ductile damage model is utilised in this work to predict the fracture initiation and propagation in uniaxial tensile testing and the SPIF process. To calculate the Hooputra parameters at different temperatures the values of fracture strain and corresponding stress triaxiality from Table 2 are plotted as five isolated points at each temperature. To obtain the best match between Eq.1 and the five isolated points, a curve fitting procedure is carried out using MATLAB tools. The best-fitting values of a, b, and c represent the parameters of the Hooputra ductile damage model. Fig. 4 depicts the Hooputra damage curves at various temperatures based on data from Table 2, and Table 3 presents the relevant Hooputra parameters. According to Fig. 4 and Table 3, the temperature affects the Hooputra damage curve and Hooputra parameters; hence, the Hooputra ductile damage model should be modified to account for temperature impacts.
<figure><img src="image_1534.png" /><figcaption>Fig. 4. Hooputra ductile damage curve at different temperatures.</figcaption></figure>

<table><caption>Table 3. Hooputra parameters for PEEK at different temperatures.</caption><thead><tr><th>Parameters</th><th>T 22 [°C]</th><th>T80 [°C]</th><th>T120 [°C]</th><th>T150 [°C]</th></tr></thead><tbody><tr><td>a</td><td>2.034</td><td>2.715</td><td>3.863</td><td>4.063</td></tr><tr><td>b</td><td>-0.1318</td><td>-0.0389</td><td>-0.0181</td><td>-0.03111</td></tr><tr><td>c</td><td>0.97</td><td>1.567</td><td>2.133</td><td>2.099</td></tr></tbody></table>

## Fracture prediction

To validate the potential of the Hooputra ductile damage model to accurately capture the PEEK fracture under different temperatures and strain states, such as uniaxial and plane strain states, the FE results of a smooth tensile sample at room temperature and 150 °C, as well as the SPIF test at 80 °C, were compared to the experimental results. Fig. 5 (a) depicts the predicted damage evolution of a smooth tensile sample at room and 150 °C temperatures. The graphic clearly shows that the damage grows faster at ambient temperature than at 150 °C. This is owing to PEEK's low formability and high stress triaxiality at room temperature as compared to 150 °C. Furthermore, when the deforming temperature is 150 °C, the displacement to fracture is greater. Fig. 5 (b) shows a comparison of the experimental and numerical tensile stress-strain diagrams at room temperature and 150 °C. The figure clearly shows that the mechanical response of PEEK changes with temperature, and fracture initiation is temperature sensitive. The FE prediction of PEEK fracture using Hooputra ductile damage model is comparable and closely matches the experimental data.

Fig. 6 depicts the damage evolution in the SPIF component under plane strain at 80 °C. It is noted that during the early stages (low forming angles) of the SPIF process, the damage grows uniformly and close to zero, and when the stretching in the forming wall reaches a particular value at wall angle 58°, the damage begins to increase rapidly, leading to fracture in the transition region between the forming wall and the base of the truncated cone at wall angle 77.6°. The numerical forming angle were compared with the experimental one to validate the prediction results of the Hooputra ductile damage model to capture the fracture in the SPIF process of hyperbolic truncated cone shape. The primary results showed that the fracture occurred early according to Hooputra ductile damage at forming angle equal to 77.6° while the experimental fracture occurred at 88.24° (see Fig. 6). The possible explanation for this, the heat generated by the friction between the forming tool and the PEEK sheet and due to the plastic deformation is not taken into consideration
<figure><img src="image_1538.png" /><figcaption>Fig. 4. Hooputra ductile damage curve at different temperatures.</figcaption></figure>

<table><caption>Table 3. Hooputra parameters for PEEK at different temperatures.</caption><thead><tr><th>Parameters</th><th>T 22 [°C]</th><th>T80 [°C]</th><th>T120 [°C]</th><th>T150 [°C]</th></tr></thead><tbody><tr><td>a</td><td>2.034</td><td>2.715</td><td>3.863</td><td>4.063</td></tr><tr><td>b</td><td>-0.1318</td><td>-0.0389</td><td>-0.0181</td><td>-0.03111</td></tr><tr><td>c</td><td>0.97</td><td>1.567</td><td>2.133</td><td>2.099</td></tr></tbody></table>

## Fracture prediction

To validate the potential of the Hooputra ductile damage model to accurately capture the PEEK fracture under different temperatures and strain states, such as uniaxial and plane strain states, the FE results of a smooth tensile sample at room temperature and 150 °C, as well as the SPIF test at 80 °C, were compared to the experimental results. Fig. 5 (a) depicts the predicted damage evolution of a smooth tensile sample at room and 150 °C temperatures. The graphic clearly shows that the damage grows faster at ambient temperature than at 150 °C. This is owing to PEEK's low formability and high stress triaxiality at room temperature as compared to 150 °C. Furthermore, when the deforming temperature is 150 °C, the displacement to fracture is greater. Fig. 5 (b) shows a comparison of the experimental and numerical tensile stress-strain diagrams at room temperature and 150 °C. The figure clearly shows that the mechanical response of PEEK changes with temperature, and fracture initiation is temperature sensitive. The FE prediction of PEEK fracture using Hooputra ductile damage model is comparable and closely matches the experimental data.

Fig. 6 depicts the damage evolution in the SPIF component under plane strain at 80 °C. It is noted that during the early stages (low forming angles) of the SPIF process, the damage grows uniformly and close to zero, and when the stretching in the forming wall reaches a particular value at wall angle 58°, the damage begins to increase rapidly, leading to fracture in the transition region between the forming wall and the base of the truncated cone at wall angle 77.6°. The numerical forming angle were compared with the experimental one to validate the prediction results of the Hooputra ductile damage model to capture the fracture in the SPIF process of hyperbolic truncated cone shape. The primary results showed that the fracture occurred early according to Hooputra ductile damage at forming angle equal to 77.6° while the experimental fracture occurred at 88.24° (see Fig. 6). The possible explanation for this, the heat generated by the friction between the forming tool and the PEEK sheet and due to the plastic deformation is not taken into consideration
in Hooputra damage model, which means the Hooputra ductile damage curve is constant in FE simulation, and according to the findings of this study, this curve changes with temperature and the safe region expands with increasing temperature. Particularly, in the SPIF FE model, the stress triaxiality-fracture strain values are compared with the Hooputra ductile damage curve, and if the value is below the curve, no damage occurs; however, if the value is above the curve, damage occurs in the FE model; however, the values of stress triaxiality-fracture strain change with increasing temperature (fracture strain increases and stress triaxiality decreases) due to the plastic deformation and friction. As a result, the Hooputra ductile damage model should be developed to account for the effect of temperature in order to be utilised to predict fracture in polymeric materials.

<figure><img src="image_1535.png" /><figcaption>Fig. 5. (a) Uniaxial damage evolution at various temperatures, and (b) comparison of numerical and experimental stress-strain curves at various temperatures.</figcaption></figure>

## Conclusions

The influence of temperature on stress triaxiality and Hooputra ductile damage curve of PEEK was addressed in this work, and the Hooputra ductile damage results of fracture prediction of uniaxial tensile test and SPIF process were validated by the experimental work. The findings of this research can be summarised as follows:

* The stress triaxiality formulas based on geometric dimensions are not sufficient for notched samples because the radius of the notched sample is no longer circular after plastic deformation.

* The Hooputra ductile damage curve of PEEK material is temperature sensitive and changes with deformation temperature.

* When the influence of temperature is addressed in Hooputra parameter calculations, the uniaxial fracture is accurately predicted using the Hooputra ductile damage model. The fracture in the SPIF process is predicted with a percentage of error; this could be because the influence of friction-generated temperature was not taken into account in the FE model and the Hooputra ductile damage curve is constant throughout the FE simulation.

* According to this investigation, the Hooputra ductile damage model should be modified to take temperature into account in order to apply this model to predict fracture initiation and propagation in polymeric materials.
in Hooputra damage model, which means the Hooputra ductile damage curve is constant in FE simulation, and according to the findings of this study, this curve changes with temperature and the safe region expands with increasing temperature. Particularly, in the SPIF FE model, the stress triaxiality-fracture strain values are compared with the Hooputra ductile damage curve, and if the value is below the curve, no damage occurs; however, if the value is above the curve, damage occurs in the FE model; however, the values of stress triaxiality-fracture strain change with increasing temperature (fracture strain increases and stress triaxiality decreases) due to the plastic deformation and friction. As a result, the Hooputra ductile damage model should be developed to account for the effect of temperature in order to be utilised to predict fracture in polymeric materials.

<figure><img src="image_1539.png" /><figcaption>Fig. 5. (a) Uniaxial damage evolution at various temperatures, and (b) comparison of numerical and experimental stress-strain curves at various temperatures.</figcaption></figure>

## Conclusions

The influence of temperature on stress triaxiality and Hooputra ductile damage curve of PEEK was addressed in this work, and the Hooputra ductile damage results of fracture prediction of uniaxial tensile test and SPIF process were validated by the experimental work. The findings of this research can be summarised as follows:

* The stress triaxiality formulas based on geometric dimensions are not sufficient for notched samples because the radius of the notched sample is no longer circular after plastic deformation.

* The Hooputra ductile damage curve of PEEK material is temperature sensitive and changes with deformation temperature.

* When the influence of temperature is addressed in Hooputra parameter calculations, the uniaxial fracture is accurately predicted using the Hooputra ductile damage model. The fracture in the SPIF process is predicted with a percentage of error; this could be because the influence of friction-generated temperature was not taken into account in the FE model and the Hooputra ductile damage curve is constant throughout the FE simulation.

* According to this investigation, the Hooputra ductile damage model should be modified to take temperature into account in order to apply this model to predict fracture initiation and propagation in polymeric materials.
<figure><img src="image_1536.png" /><figcaption>Fig. 6. Plane strain damage evolution at 80 °C.</figcaption></figure>

**Acknowledgements**

This work was supported by the UK Engineering and Physical Science Research Council [Grant No. EP/W010089/1].

**References**

[1] Zheng, Bing, Haitao Wang, Zhigao Huang, Yi Zhang, Huamin Zhou, and Dequn Li, Experimental investigation and constitutive modeling of the deformation behavior of Poly-Ether-Ether-Ketone at elevated temperatures, Polymer Testing. 63 (2017) 349-359. https://doi.org/10.1016/j.polymertesting.2017.08.020

[2] Chen, Lin-feng, Fei Chen, Shakir Gatea, and Hengan Ou, PEEK based cranial reconstruction using thermal assisted incremental sheet forming, Proc IMechE Part B: J Engineering Manufacture. 236 (6-7) (2022) 997-1004. https://doi.org/10.1177/09544054211045904

[3] Chen, Fei, Shakir Gatea, Hengan Ou, Bin Lu, and Hui Long, Fracture characteristics of PEEK at various stress triaxialities, Journal of the mechanical behavior of biomedical materials. 64 (2016) 173-186. https://doi.org/10.1016/j.jmbbm.2016.07.027

[4] Johnson, Gordon R and William H Cook, Fracture characteristics of three metals subjected to various strains, strain rates, temperatures and pressures, Engineering fracture mechanics. 21 (1) (1985) 31-48. https://doi.org/10.1016/0013-7944(85)90052-9

[5] Kanvinde, AM and GG Deierlein, The void growth model and the stress modified critical strain model to predict ductile fracture in structural steels, Journal of Structural Engineering. 132 (12) (2006) 1907-1918. https://doi.org/10.1061/(ASCE)0733-9445(2006)132:12(1907)
<figure><img src="image_1540.png" /><figcaption>Fig. 6. Plane strain damage evolution at 80 °C.</figcaption></figure>

**Acknowledgements**

This work was supported by the UK Engineering and Physical Science Research Council [Grant No. EP/W010089/1].

**References**

[1] Zheng, Bing, Haitao Wang, Zhigao Huang, Yi Zhang, Huamin Zhou, and Dequn Li, Experimental investigation and constitutive modeling of the deformation behavior of Poly-Ether-Ether-Ketone at elevated temperatures, Polymer Testing. 63 (2017) 349-359. https://doi.org/10.1016/j.polymertesting.2017.08.020

[2] Chen, Lin-feng, Fei Chen, Shakir Gatea, and Hengan Ou, PEEK based cranial reconstruction using thermal assisted incremental sheet forming, Proc IMechE Part B: J Engineering Manufacture. 236 (6-7) (2022) 997-1004. https://doi.org/10.1177/09544054211045904

[3] Chen, Fei, Shakir Gatea, Hengan Ou, Bin Lu, and Hui Long, Fracture characteristics of PEEK at various stress triaxialities, Journal of the mechanical behavior of biomedical materials. 64 (2016) 173-186. https://doi.org/10.1016/j.jmbbm.2016.07.027

[4] Johnson, Gordon R and William H Cook, Fracture characteristics of three metals subjected to various strains, strain rates, temperatures and pressures, Engineering fracture mechanics. 21 (1) (1985) 31-48. https://doi.org/10.1016/0013-7944(85)90052-9

[5] Kanvinde, AM and GG Deierlein, The void growth model and the stress modified critical strain model to predict ductile fracture in structural steels, Journal of Structural Engineering. 132 (12) (2006) 1907-1918. https://doi.org/10.1061/(ASCE)0733-9445(2006)132:12(1907)
[6] Brünig, Michael, Oliver Chyra, Daniel Albrecht, Larissa Driemeier, and Marcílio Alves, A ductile damage criterion at various stress triaxialities, International journal of plasticity. 24 (10) (2008) 1731-1755. https://doi.org/10.1016/j.ijplas.2007.12.001

[7] Peng, Zengli, Haisheng Zhao, and Xin Li, New ductile fracture model for fracture prediction ranging from negative to high stress triaxiality, International Journal of Plasticity. 145 (2021) 103057. https://doi.org/10.1016/j.ijplas.2021.103057

[8] Hooputra, H, H Gese, H Dell, and H Werner, A comprehensive failure model for crashworthiness simulation of aluminium extrusions, International Journal of Crashworthiness. 9 (5) (2004) 449-464. https://doi.org/10.1533/ijcr.2004.0289

[9] Wierzbicki, Tomasz, Yingbin Bao, Young-Woong Lee, and Yuanli Bai, Calibration and evaluation of seven fracture models, International Journal of Mechanical Sciences. 47 (4-5) (2005) 719-743. https://doi.org/10.1016/j.ijmecsci.2005.03.003

[10] Zadpoor, Amir Abbas, Jos Sinke, and Rinze Benedictus, Formability prediction of high strength aluminum sheets, International Journal of Plasticity, 2009. 25(12): p. 2269-2297. https://doi.org/10.1016/j.ijplas.2009.02.005

[11] Haji Aboutalebi, F and A Banihashemi, Numerical estimation and practical validation of Hooputra's ductile damage parameters, The International Journal of Advanced Manufacturing Technology. 75 (2014) 1701-1710. https://doi.org/10.1007/s00170-014-6275-8

[12] Estrada, Quirino, Dariusz Szwedowicz, Jesús Silva-Aceves, Tadeusz Majewski, Julio Vergara-Vazquez, and Alejandro Rodriguez-Mendez, Crashworthiness behavior of aluminum profiles with holes considering damage criteria and damage evolution, International Journal of Mechanical Sciences. 131 (2017) 776-791. https://doi.org/10.1016/j.ijmecsci.2017.07.042

[13] Zhu, Yazhi and Michael D Engelhardt, Prediction of ductile fracture for metal alloys using a shear modified void growth model, Engineering Fracture Mechanics. 190 (2018) 491-513. https://doi.org/10.1016/j.engfracmech.2017.12.042

[14] Zhu, Hui, Hengan Ou, and Atanas Popov, Incremental sheet forming of thermoplastics: a review, The International Journal of Advanced Manufacturing Technology. 111 (2020) 565-587. https://doi.org/10.1007/s00170-020-06056-5

[15] Yang, Zhiyuan, Fei Chen, Shakir Gatea, and Hengan Ou, Design of the novel hot incremental sheet forming experimental setup, characterization of formability behavior of polyether-ether-ketone (PEEK), The International Journal of Advanced Manufacturing Technology. 106 (2020) 5365-5381. https://doi.org/10.1007/s00170-020-05035-0

[16] Yang, Zhiyun and Fei Chen, Mechanism of twist in incremental sheet forming of thermoplastic polymer, Materials & Design. 195 (2020) 108997. https://doi.org/10.1016/j.matdes.2020.108997

[17] Rosa-Sainz, Ana, María Luisa García-Romeu, I Ferrer, María Beatriz Silva, and G Centeno, On the effective peek application for customized cranio-maxillofacial prostheses: An experimental formability analysis, Journal of Manufacturing Processes. 86 (2023) 66-84. https://doi.org/10.1016/j.jmapro.2022.12.044

[18] Bai, Yuanli, Xiaoqing Teng, and Tomasz Wierzbicki, On the application of stress triaxiality formula for plane strain fracture testing. J Eng MaterTechnol. 131 (2) (2009) 021002. https://doi.org/10.1115/1.3078390
[6] Brünig, Michael, Oliver Chyra, Daniel Albrecht, Larissa Driemeier, and Marcílio Alves, A ductile damage criterion at various stress triaxialities, International journal of plasticity. 24 (10) (2008) 1731-1755. https://doi.org/10.1016/j.ijplas.2007.12.001

[7] Peng, Zengli, Haisheng Zhao, and Xin Li, New ductile fracture model for fracture prediction ranging from negative to high stress triaxiality, International Journal of Plasticity. 145 (2021) 103057. https://doi.org/10.1016/j.ijplas.2021.103057

[8] Hooputra, H, H Gese, H Dell, and H Werner, A comprehensive failure model for crashworthiness simulation of aluminium extrusions, International Journal of Crashworthiness. 9 (5) (2004) 449-464. https://doi.org/10.1533/ijcr.2004.0289

[9] Wierzbicki, Tomasz, Yingbin Bao, Young-Woong Lee, and Yuanli Bai, Calibration and evaluation of seven fracture models, International Journal of Mechanical Sciences. 47 (4-5) (2005) 719-743. https://doi.org/10.1016/j.ijmecsci.2005.03.003

[10] Zadpoor, Amir Abbas, Jos Sinke, and Rinze Benedictus, Formability prediction of high strength aluminum sheets, International Journal of Plasticity, 2009. 25(12): p. 2269-2297. https://doi.org/10.1016/j.ijplas.2009.02.005

[11] Haji Aboutalebi, F and A Banihashemi, Numerical estimation and practical validation of Hooputra's ductile damage parameters, The International Journal of Advanced Manufacturing Technology. 75 (2014) 1701-1710. https://doi.org/10.1007/s00170-014-6275-8

[12] Estrada, Quirino, Dariusz Szwedowicz, Jesús Silva-Aceves, Tadeusz Majewski, Julio Vergara-Vazquez, and Alejandro Rodriguez-Mendez, Crashworthiness behavior of aluminum profiles with holes considering damage criteria and damage evolution, International Journal of Mechanical Sciences. 131 (2017) 776-791. https://doi.org/10.1016/j.ijmecsci.2017.07.042

[13] Zhu, Yazhi and Michael D Engelhardt, Prediction of ductile fracture for metal alloys using a shear modified void growth model, Engineering Fracture Mechanics. 190 (2018) 491-513. https://doi.org/10.1016/j.engfracmech.2017.12.042

[14] Zhu, Hui, Hengan Ou, and Atanas Popov, Incremental sheet forming of thermoplastics: a review, The International Journal of Advanced Manufacturing Technology. 111 (2020) 565-587. https://doi.org/10.1007/s00170-020-06056-5

[15] Yang, Zhiyuan, Fei Chen, Shakir Gatea, and Hengan Ou, Design of the novel hot incremental sheet forming experimental setup, characterization of formability behavior of polyether-ether-ketone (PEEK), The International Journal of Advanced Manufacturing Technology. 106 (2020) 5365-5381. https://doi.org/10.1007/s00170-020-05035-0

[16] Yang, Zhiyun and Fei Chen, Mechanism of twist in incremental sheet forming of thermoplastic polymer, Materials & Design. 195 (2020) 108997. https://doi.org/10.1016/j.matdes.2020.108997

[17] Rosa-Sainz, Ana, María Luisa García-Romeu, I Ferrer, María Beatriz Silva, and G Centeno, On the effective peek application for customized cranio-maxillofacial prostheses: An experimental formability analysis, Journal of Manufacturing Processes. 86 (2023) 66-84. https://doi.org/10.1016/j.jmapro.2022.12.044

[18] Bai, Yuanli, Xiaoqing Teng, and Tomasz Wierzbicki, On the application of stress triaxiality formula for plane strain fracture testing. J Eng MaterTechnol. 131 (2) (2009) 021002. https://doi.org/10.1115/1.3078390
# Spindle speed effect on the ISF processing of materials with different thermal conductivities

QIN Zhenyuan<fnref n="1" />,ª, GATEA Shakir<fnref n="1" />,<fnref n="2" />,ᵇ, and OU Hengan<fnref n="1" />,ᶜ*

**Keywords:** Incremental Sheet Forming, Thermal Conductivity, Springback, Surface Finish, Finite Element Method

**Abstract.** Incremental sheet forming (ISF) is an advanced forming technology with high flexibility and excellent adaptability. However, conventional ISF encounters challenges in processing of hard-to-deformed materials at room temperature. To address this problem, friction stir incremental forming (FSIF) has been introduced to improve material formability at high temperatures. The aim of this work is to investigate the spindle speed effect on the forming quality for materials with different thermal conductivities by using an experimental and FE simulation. The experimental results presented the combined effect of spindle speed and conductivity on the springback and surface finish of AA1050 and CP Ti Grade 1. A good agreement was obtained in the temperature increase and distribution from experimental testing and FE simulation.

**Introduction**

Incremental sheet forming (ISF), also known as single point incremental forming (SPIF), is characterised as a flexible sheet forming technology due to the fact that it does not require specific dies and is capable of producing components with complex shapes [1,2]. Although conventional ISF possesses some advantages as compared to other sheet metal forming processes, it is still a technical challenge to use ISF to form hard-to-deform materials. To overcome this limitation, friction stir incremental forming (FSIF) has been developed to improve the formability of the hard-to-deform materials through friction heat induced by the rotating tool [3].

Due to its high applicability and simplicity, FSIF has attracted considerable research attention in the past decade. Wang et al. [4] demonstrated the formability of AA5052-H32 and AA2024-T can be significantly improved accompanied with reasonable surface quality when applying a suitable range of rotation speed. A work by Grün et al. [5] highlighted the influence of rotation speed on the forming limit of Ti6Al4V alloy. Formisano et al. [6] revealed the effect of processing parameters on the forming temperature in FSIF by conducting experiments on the polycarbonate (PC) sheets.

Besides the above experimental observations, some numerical studies on FSIF have been also reported by researchers. In the numerical work of Li and Wang [7], the non-rotating tool with equivalent temperature was adopted to enable that the same temperature variation trend as the actual testing case with rotating tool can be achieved in highly reduced simulation time. Cai et al. [8] performed FE simulations to explore the effects of process variables, i.e., feed rate, spindle speed, wall angle and vertical step size, on temperature evolution, thickness distribution, and central bulge. The simulation results were in good agreement with the experimental results. Nguyen et al. [9] used commercial FE software Abaqus incorporating VUMAT subroutine code to evaluate the fracture behaviour of AZ31 magnesium alloy in FSIF. It was shown that the fracture

<figure><img src="image_1537.png" /></figure>

<footnote n="1">¹Department of Mechanical, Materials and Manufacturing Engineering, Faculty of Engineering, The University of Nottingham, Nottingham NG7 2RD, UK</footnote>

<footnote n="2">²Department of Materials Engineering, Faculty of Engineering, University of Kufa, Al-Najaf, Iraq</footnote>

<footnote>ªeaxzq1@nottingham.ac.uk, ᵇezzsmg@exmail.nottingham.ac.uk, ᶜh.ou@nottingham.ac.uk</footnote>
# Spindle speed effect on the ISF processing of materials with different thermal conductivities

QIN Zhenyuan<sup>1,a</sup>, GATEA Shakir<sup>1,2,b</sup>, and OU Hengan<sup>1,c*</sup>

<sup>1</sup>Department of Mechanical, Materials and Manufacturing Engineering, Faculty of Engineering, The University of Nottingham, Nottingham NG7 2RD, UK

<sup>2</sup>Department of Materials Engineering, Faculty of Engineering, University of Kufa, Al-Najaf, Iraq

<sup>a</sup>eaxzq1@nottingham.ac.uk, <sup>b</sup>ezzsmg@exmail.nottingham.ac.uk, <sup>c</sup>h.ou@nottingham.ac.uk

**Keywords:** Incremental Sheet Forming, Thermal Conductivity, Springback, Surface Finish, Finite Element Method

**Abstract.** Incremental sheet forming (ISF) is an advanced forming technology with high flexibility and excellent adaptability. However, conventional ISF encounters challenges in processing of hard-to-deformed materials at room temperature. To address this problem, friction stir incremental forming (FSIF) has been introduced to improve material formability at high temperatures. The aim of this work is to investigate the spindle speed effect on the forming quality for materials with different thermal conductivities by using an experimental and FE simulation. The experimental results presented the combined effect of spindle speed and conductivity on the springback and surface finish of AA1050 and CP Ti Grade 1. A good agreement was obtained in the temperature increase and distribution from experimental testing and FE simulation.

**Introduction**

Incremental sheet forming (ISF), also known as single point incremental forming (SPIF), is characterised as a flexible sheet forming technology due to the fact that it does not require specific dies and is capable of producing components with complex shapes [1,2]. Although conventional ISF possesses some advantages as compared to other sheet metal forming processes, it is still a technical challenge to use ISF to form hard-to-deform materials. To overcome this limitation, friction stir incremental forming (FSIF) has been developed to improve the formability of the hard-to-deform materials through friction heat induced by the rotating tool [3].

Due to its high applicability and simplicity, FSIF has attracted considerable research attention in the past decade. Wang et al. [4] demonstrated the formability of AA5052-H32 and AA2024-T can be significantly improved accompanied with reasonable surface quality when applying a suitable range of rotation speed. A work by Grün et al. [5] highlighted the influence of rotation speed on the forming limit of Ti6Al4V alloy. Formisano et al. [6] revealed the effect of processing parameters on the forming temperature in FSIF by conducting experiments on the polycarbonate (PC) sheets.

Besides the above experimental observations, some numerical studies on FSIF have been also reported by researchers. In the numerical work of Li and Wang [7], the non-rotating tool with equivalent temperature was adopted to enable that the same temperature variation trend as the actual testing case with rotating tool can be achieved in highly reduced simulation time. Cai et al. [8] performed FE simulations to explore the effects of process variables, i.e., feed rate, spindle speed, wall angle and vertical step size, on temperature evolution, thickness distribution, and central bulge. The simulation results were in good agreement with the experimental results. Nguyen et al. [9] used commercial FE software Abaqus incorporating VUMAT subroutine code to evaluate the fracture behaviour of AZ31 magnesium alloy in FSIF. It was shown that the fracture

<figure><img src="image_1541.png" /></figure>
depth under specific forming condition was successfully predicted by the proposed numerical model.

Although aforementioned literature provides a profound understanding about FSIF in different aspects, no attempts have been made to evaluate the combined effect of spindle speed and thermal conductivities on the quality of FSIF processed parts. Therefore, this work aims to investigate the thermal response in FSIF processing of sheet materials with different thermal properties and to provide an insight into the interrelationship between the forming quality and thermal properties of the formed parts based on experimental testing and FE simulation.

## Experimental testing

In this work, the AA1050 aluminium alloy and CP Ti Grade 1 pure titanium were used as workpiece, with the dimension of 150×150×0.7mm and 150×150×0.5mm, respectively. Both materials were formed by using a high-speed steel tool with a diameter of 10mm. The physical and thermal properties of workpiece and forming tool are presented in Table1.

<table><caption>Table 1. The physical and thermal properties of workpiece and forming tool.</caption><thead><tr><th>Material</th><th>Density [kg/m³]</th><th>Poisson's ratio</th><th>Young's modulus [GPa]</th><th>Yield stress [MPa]</th><th>Specific heat[J/(kg.K)]</th><th>Conductivity [W/(m.K)]</th></tr></thead><tbody><tr><td>AA1050</td><td>2700</td><td>0.33</td><td>71</td><td>90</td><td>900</td><td>231</td></tr><tr><td>CP Ti Grade 1</td><td>4540</td><td>0.34</td><td>105</td><td>221</td><td>520</td><td>16.4</td></tr><tr><td>High-speed steel</td><td>7600</td><td>0.27</td><td>200</td><td>1000</td><td>460</td><td>41</td></tr></tbody></table>

The ISF was carried out on a CNC milling machine, a hyperbolic truncated cone with continuously varying wall angles from 30° to 90° was formed by using a hemispherical tool with a diameter of 10mm. Two spindle speeds (0 rpm, 2500 rpm) were used in the test, while other process parameters remained unchanged: with feed rate of 1500 mm/min, step size of 0.5 mm, and with use of Rando HD68 lubricant. The forming forces were measured by a load cell mounted below the fixture. The temperature variation of specific tracking positions with equal spacing of 6mm (see Fig. 1) was monitored by a series of thermocouples.
depth under specific forming condition was successfully predicted by the proposed numerical model.

Although aforementioned literature provides a profound understanding about FSIF in different aspects, no attempts have been made to evaluate the combined effect of spindle speed and thermal conductivities on the quality of FSIF processed parts. Therefore, this work aims to investigate the thermal response in FSIF processing of sheet materials with different thermal properties and to provide an insight into the interrelationship between the forming quality and thermal properties of the formed parts based on experimental testing and FE simulation.

## Experimental testing

In this work, the AA1050 aluminium alloy and CP Ti Grade 1 pure titanium were used as workpiece, with the dimension of 150×150×0.7mm and 150×150×0.5mm, respectively. Both materials were formed by using a high-speed steel tool with a diameter of 10mm. The physical and thermal properties of workpiece and forming tool are presented in Table1.

<table><caption>Table 1. The physical and thermal properties of workpiece and forming tool.</caption><thead><tr><th>Material</th><th>Density [kg/m³]</th><th>Poisson's ratio</th><th>Young's modulus [GPa]</th><th>Yield stress [MPa]</th><th>Specific heat[J/(kg.K)]</th><th>Conductivity [W/(m.K)]</th></tr></thead><tbody><tr><td>AA1050</td><td>2700</td><td>0.33</td><td>71</td><td>90</td><td>900</td><td>231</td></tr><tr><td>CP Ti Grade 1</td><td>4540</td><td>0.34</td><td>105</td><td>221</td><td>520</td><td>16.4</td></tr><tr><td>High-speed steel</td><td>7600</td><td>0.27</td><td>200</td><td>1000</td><td>460</td><td>41</td></tr></tbody></table>

The ISF was carried out on a CNC milling machine, a hyperbolic truncated cone with continuously varying wall angles from 30° to 90° was formed by using a hemispherical tool with a diameter of 10mm. Two spindle speeds (0 rpm, 2500 rpm) were used in the test, while other process parameters remained unchanged: with feed rate of 1500 mm/min, step size of 0.5 mm, and with use of Rando HD68 lubricant. The forming forces were measured by a load cell mounted below the fixture. The temperature variation of specific tracking positions with equal spacing of 6mm (see Fig. 1) was monitored by a series of thermocouples.
<figure><img src="image_1538.png" /><figcaption>Fig. 1. FSIF tool rotation, movement and thermocouple tracking positions (unit: mm).</figcaption></figure>

Finite element simulation

A coupled thermo-mechanical FE model was built in Abaqus/explicit to simulate the ISF process under specific processing condition. To obtain accurate solution to three- dimensional heat transfer problem, both workpiece and forming tool were applied with the 3D thermally coupled solid brick elements with eight nodes (C3D8T). In FE model, the workpiece was simulated as a deformable body, while the forming tool was supposed as a rigid body and its movements (along X, Y and Z directions) were constrained at a reference point. In addition, the material properties shown in Table 1 were used in FE as input values. The contact behavior at tool-workpiece interface was defined by using embedded penalty contact algorithm. By referring to previous research work, the friction coefficients were set to 0.09 and 0.167 for AA1050 and CP Ti Grade 1, respectively. The heat transfer coefficient was set to 20 W/ (m².K). The simulation results including forming depth and forming temperature were visualized and analysed in the output module.

Results and discussion

The forming force and forming temperature

Fig. 2 presents the variation of three components of forming forces including vertical force (Fz) and horizontal force (Fx and Fy) at different spindle speeds. For AA1050 and CP Ti Grade 1, Fz increased gradually until the peak force reached, and then monotonically decreased but exhibited different rate of decline. This can be ascribed to different flow and thinning behaviour of both materials. The variation of Fx and Fy followed a sinusoidal trend with increasing amplitude. When the workpiece was formed contour by contour, Fx and Fy shown similar patterns but with a phase difference of π/2. With increase of spindle speed, the vertical and horizontal forces displayed a drop for both materials. This is because the spindle speed has significant contribution to the friction heat, resulting in material softening effect and reduction of yield stress.
<figure><img src="image_1542.png" /><figcaption>Fig. 1. FSIF tool rotation, movement and thermocouple tracking positions (unit: mm).</figcaption></figure>

Finite element simulation

A coupled thermo-mechanical FE model was built in Abaqus/explicit to simulate the ISF process under specific processing condition. To obtain accurate solution to three- dimensional heat transfer problem, both workpiece and forming tool were applied with the 3D thermally coupled solid brick elements with eight nodes (C3D8T). In FE model, the workpiece was simulated as a deformable body, while the forming tool was supposed as a rigid body and its movements (along X, Y and Z directions) were constrained at a reference point. In addition, the material properties shown in Table 1 were used in FE as input values. The contact behavior at tool-workpiece interface was defined by using embedded penalty contact algorithm. By referring to previous research work, the friction coefficients were set to 0.09 and 0.167 for AA1050 and CP Ti Grade 1, respectively. The heat transfer coefficient was set to 20 W/ (m².K). The simulation results including forming depth and forming temperature were visualized and analysed in the output module.

Results and discussion

The forming force and forming temperature

Fig. 2 presents the variation of three components of forming forces including vertical force (Fz) and horizontal force (Fx and Fy) at different spindle speeds. For AA1050 and CP Ti Grade 1, Fz increased gradually until the peak force reached, and then monotonically decreased but exhibited different rate of decline. This can be ascribed to different flow and thinning behaviour of both materials. The variation of Fx and Fy followed a sinusoidal trend with increasing amplitude. When the workpiece was formed contour by contour, Fx and Fy shown similar patterns but with a phase difference of π/2. With increase of spindle speed, the vertical and horizontal forces displayed a drop for both materials. This is because the spindle speed has significant contribution to the friction heat, resulting in material softening effect and reduction of yield stress.
<figure><img src="image_1539.png" /><figcaption>Fig. 2. The variation of forming forces for (a) AA1050 at 0 rpm (b) AA1050 at 2500 rpm (c) CP Ti Grade 1 at 0 rpm and (d) CP Ti Grade 1 at 2500 rpm.</figcaption></figure>

During the testing, the temperature variation of specific tracking positions was measured by thermocouples. If these measured results are agreed with the temperature values given by FE model, the maximum forming temperature can be obtained from simulation results. The peak temperatures were measured at P4 and P3 (tracking positions shown in Fig. 1) for AA1050 and CP Ti Grade 1, respectively, after P4 and P3, the forming temperatures of the two materials reached the steady state. As can be seen in Fig. 3, as rotation speed increased from 0 rpm to 2500 rpm, the maximum forming temperature elevated from 24.6 °C to 97.45 °C for AA1050, and from 36.2 °C to 247.2 °C for CP Ti Grade 1, respectively. This dissimilar thermal response can be mainly attributed to two reasons: (i) The maximum forming force of CP Ti Grade 1 is greater than that of AA1050, leading to larger amount of friction induced heat generation. (ii) Since the thermal diffusivity and conductivity of CP Ti Grade 1 is much lower than that of AA1050, the heat flux tends to be slowly transferred from local forming zone to the surrounding regions, thereby causing obvious thermal concentration and higher temperature rise around tool-workpiece contact area. Giuseppina et al. [10] also reported similar observation and captured the effect of thermal properties on the forming temperature by performing ISF testing on AA5754 and Ti6Al4V.
<figure><img src="image_1543.png" /><figcaption>Fig. 2. The variation of forming forces for (a) AA1050 at 0 rpm (b) AA1050 at 2500 rpm (c) CP Ti Grade 1 at 0 rpm and (d) CP Ti Grade 1 at 2500 rpm.</figcaption></figure>

During the testing, the temperature variation of specific tracking positions was measured by thermocouples. If these measured results are agreed with the temperature values given by FE model, the maximum forming temperature can be obtained from simulation results. The peak temperatures were measured at P4 and P3 (tracking positions shown in Fig. 1) for AA1050 and CP Ti Grade 1, respectively, after P4 and P3, the forming temperatures of the two materials reached the steady state. As can be seen in Fig. 3, as rotation speed increased from 0 rpm to 2500 rpm, the maximum forming temperature elevated from 24.6 °C to 97.45 °C for AA1050, and from 36.2 °C to 247.2 °C for CP Ti Grade 1, respectively. This dissimilar thermal response can be mainly attributed to two reasons: (i) The maximum forming force of CP Ti Grade 1 is greater than that of AA1050, leading to larger amount of friction induced heat generation. (ii) Since the thermal diffusivity and conductivity of CP Ti Grade 1 is much lower than that of AA1050, the heat flux tends to be slowly transferred from local forming zone to the surrounding regions, thereby causing obvious thermal concentration and higher temperature rise around tool-workpiece contact area. Giuseppina et al. [10] also reported similar observation and captured the effect of thermal properties on the forming temperature by performing ISF testing on AA5754 and Ti6Al4V.
<figure><img src="image_1540.png" /><figcaption>Fig. 3. The maximum forming temperature of AA1050 and CP Ti Grade 1 at different spindle speeds.</figcaption></figure>

## Combined effect of spindle speed and conductivity on the springback and surface finish

As discussed in above section, although under the same forming condition, a distinct difference of temperature variation was noted during processing of AA1050 and CP Ti Grade 1. This is because two determining factors (spindle speed and conductivity) are involved concurrently. Moreover, considering that the thermal effect has significant influence on the forming quality, it is necessary to study the combined effect of spindle speed and conductivity on the springback and surface finish.

<figure><img src="image_1541.png" /><figcaption>Fig. 4. The geometry of formed part for (a) AA1050 at 0 rpm (b) AA1050 at 2500 rpm (c) CP Ti Grade 1 at 0 rpm and (d) CP Ti Grade 1 at 2500 rpm.</figcaption></figure>
<figure><img src="image_1544.png" /><figcaption>Fig. 3. The maximum forming temperature of AA1050 and CP Ti Grade 1 at different spindle speeds.</figcaption></figure>

## Combined effect of spindle speed and conductivity on the springback and surface finish

As discussed in above section, although under the same forming condition, a distinct difference of temperature variation was noted during processing of AA1050 and CP Ti Grade 1. This is because two determining factors (spindle speed and conductivity) are involved concurrently. Moreover, considering that the thermal effect has significant influence on the forming quality, it is necessary to study the combined effect of spindle speed and conductivity on the springback and surface finish.

<figure><img src="image_1545.png" /><figcaption>Fig. 4. The geometry of formed part for (a) AA1050 at 0 rpm (b) AA1050 at 2500 rpm (c) CP Ti Grade 1 at 0 rpm and (d) CP Ti Grade 1 at 2500 rpm.</figcaption></figure>
Fig. 4 shows the geometry of formed part for AA1050 and CP Ti Grade 1 under different processing conditions. It was noted that there was almost no springback for AA1050 at either 0 rpm or 2500 rpm, while the geometry accuracy was reasonable at 0 rpm but obvious springback occurred at 2500 rpm for CP Ti Grade 1. This is due to the fact that the thermal concentration creates high temperature difference between tool-workpiece contact area and other regions, leading to non-uniform heating induced springback for CP Ti Grade 1 at 2500 rpm. This phenomenon was also observed by Li et al. [11] through springback analysis on titanium tube. In contrast, high conductivity makes the heat flux more uniformly diffused throughout AA1050 workpiece to enable that the good geometry accuracy can be achieved at 2500 rpm.

<figure><img src="image_1542.png" /><figcaption>Fig. 5. The surface finish of formed part for (a) AA1050 at 0 rpm (b) AA1050 at 2500 rpm (c) CP Ti Grade 1 at 0 rpm and (d) CP Ti Grade 1 at 2500 rpm.</figcaption></figure>

As can be seen from Fig. 5, under the spindle speed of 0 rpm, both AA1050 and CP Ti Grade 1 achieved acceptable surface finish; at 2500 rpm, AA1050 still had good surface finish while clear surface degradation happened for CP Ti Grade 1. This can be explained by the following reasons: firstly, the maximum shear stress of CP Ti Grade 1 is much larger than that of AA1050, which poses a negative effect on the surface quality especially at high spindle speeds. Secondly, the thermal concentration can not only locally soften the material at tool-workpiece interface, but also introduce instantaneous spark (shown in Fig. 5(d)) during the process, thereby significantly accelerating the surface degradation.

To explain the above experimental observations from a numerical perspective, FE thermo-mechanical simulations were conducted in this work. The FE models with spindle speed of 2500 rpm for both materials are shown in Fig. 6.
Fig. 4 shows the geometry of formed part for AA1050 and CP Ti Grade 1 under different processing conditions. It was noted that there was almost no springback for AA1050 at either 0 rpm or 2500 rpm, while the geometry accuracy was reasonable at 0 rpm but obvious springback occurred at 2500 rpm for CP Ti Grade 1. This is due to the fact that the thermal concentration creates high temperature difference between tool-workpiece contact area and other regions, leading to non-uniform heating induced springback for CP Ti Grade 1 at 2500 rpm. This phenomenon was also observed by Li et al. [11] through springback analysis on titanium tube. In contrast, high conductivity makes the heat flux more uniformly diffused throughout AA1050 workpiece to enable that the good geometry accuracy can be achieved at 2500 rpm.

<figure><img src="image_1546.png" /><figcaption>Fig. 5. The surface finish of formed part for (a) AA1050 at 0 rpm (b) AA1050 at 2500 rpm (c) CP Ti Grade 1 at 0 rpm and (d) CP Ti Grade 1 at 2500 rpm.</figcaption></figure>

As can be seen from Fig. 5, under the spindle speed of 0 rpm, both AA1050 and CP Ti Grade 1 achieved acceptable surface finish; at 2500 rpm, AA1050 still had good surface finish while clear surface degradation happened for CP Ti Grade 1. This can be explained by the following reasons: firstly, the maximum shear stress of CP Ti Grade 1 is much larger than that of AA1050, which poses a negative effect on the surface quality especially at high spindle speeds. Secondly, the thermal concentration can not only locally soften the material at tool-workpiece interface, but also introduce instantaneous spark (shown in Fig. 5(d)) during the process, thereby significantly accelerating the surface degradation.

To explain the above experimental observations from a numerical perspective, FE thermo-mechanical simulations were conducted in this work. The FE models with spindle speed of 2500 rpm for both materials are shown in Fig. 6.
<figure><img src="image_1543.png" /><figcaption>Fig. 6. The FE models with spindle speed of 2500 rpm for (a) AA1050 and (b) CP Ti Grade 1.</figcaption></figure>

<figure><img src="image_1544.png" /><figcaption>Fig. 7. The temperature distribution in the final stage of process with spindle speed of 2500 rpm for (a) AA1050 and (b) CP Ti Grade 1.</figcaption></figure>

Fig. 7 presents the temperature distribution in the final stage of process with spindle speed of 2500 rpm for both materials. It is evident that the temperature is distributed in an arc shape along the toolpath. The forming temperature reached the peak value at centre of tool-workpiece interface and exhibited a regular gradient from interface to surrounding regions. The maximum temperatures obtained from FE simulation were 97.45 °C and 247.2 °C for AA1050 and CP Ti Grade 1, respectively, which were slightly larger than the measured by thermocouples (95.1 °C for AA1050 and 239.7 °C for CP Ti Grade 1). Although the maximum temperature of CP Ti Grade 1 was much higher than that of AA1050, the centre and outer edges of CP Ti Grade 1 workpiece were subjected to insufficient thermal effect, leading to significant temperature difference and thermal concentration. The simulation results captured the combined effect of spindle speed and thermal conductivities on the temperature distribution.

## Conclusions

This work presents an experimental and numerical investigation on the springback and surface finish of different materials in FSIF process, the main findings are outlined as follows:

(1) With an increase in spindle speed, the forming temperature increases and thus results in decrease of forming forces for both materials.
<figure><img src="image_1547.png" /><figcaption>Fig. 6. The FE models with spindle speed of 2500 rpm for (a) AA1050 and (b) CP Ti Grade 1.</figcaption></figure>

<figure><img src="image_1548.png" /><figcaption>Fig. 7. The temperature distribution in the final stage of process with spindle speed of 2500 rpm for (a) AA1050 and (b) CP Ti Grade 1.</figcaption></figure>

Fig. 7 presents the temperature distribution in the final stage of process with spindle speed of 2500 rpm for both materials. It is evident that the temperature is distributed in an arc shape along the toolpath. The forming temperature reached the peak value at centre of tool-workpiece interface and exhibited a regular gradient from interface to surrounding regions. The maximum temperatures obtained from FE simulation were 97.45 °C and 247.2 °C for AA1050 and CP Ti Grade 1, respectively, which were slightly larger than the measured by thermocouples (95.1 °C for AA1050 and 239.7 °C for CP Ti Grade 1). Although the maximum temperature of CP Ti Grade 1 was much higher than that of AA1050, the centre and outer edges of CP Ti Grade 1 workpiece were subjected to insufficient thermal effect, leading to significant temperature difference and thermal concentration. The simulation results captured the combined effect of spindle speed and thermal conductivities on the temperature distribution.

## Conclusions

This work presents an experimental and numerical investigation on the springback and surface finish of different materials in FSIF process, the main findings are outlined as follows:

(1) With an increase in spindle speed, the forming temperature increases and thus results in decrease of forming forces for both materials.
(2) Although under same processing condition, the surface finish and pattern of springback of the two materials are quite different due to the combined effect of spindle speed and thermal conductivities.

(3) Based on the results, it can be concluded that for material with high conductivity, the FSIF could be a viable option to achieve enhanced formability; whilst for sheet material with low conductivity, it would be better to use a different approach such as global heating (e.g., hot air heating) to obtain similar degree of formability improvement.

## Acknowledgements

This work was supported by the UK Engineering and Physical Science Research Council [Grant No. EP/W010089/1].

## References

[1] H. Zhu, H. Ou, A. Popov, Incremental sheet forming of thermoplastics: a review, The International Journal of Advanced Manufacturing Technology. 111 (2020) 565-587.
https://doi.org/10.1007/s00170-020-06056-5

[2] W. Peng, H. Ou, Deformation mechanisms and fracture in tension under cyclic bending plus compression, single point and double-sided incremental sheet forming processes, International Journal of Machine Tools and Manufacture. 184 (2023) 103980.
https://doi.org/10.1016/j.ijmachtools.2022.103980

[3] M. Li, R. Wu, S. Cai, Z. Chang, Z. Wang, J. Chen, Experimental investigation on friction-stir-assisted incremental forming with synchronous bonding of aluminum alloy and steel sheets, Journal of Materials Engineering and Performance. 29 (2020) 750-759.
https://doi.org/10.1007/s11665-020-04600-8

[4] Z. Wang, S. Cai, J. Chen, Experimental investigations on friction stir assisted single point incremental forming of low-ductility aluminum alloy sheet for higher formability with reasonable surface quality, Journal of Materials Processing Technology. 277 (2020) 116488.
https://doi.org/10.1016/j.jmatprotec.2019.116488

[5] P.A. Grün, E.H. Uheida, L. Lachmann, D. Dimitrov, G.A. Oosthuizen, Formability of titanium alloy sheets by friction stir incremental forming, The International Journal of Advanced Manufacturing Technology. 99 (2018) 1993-2003. https://doi.org/10.1007/s00170-018-2541-5

[6] A. Formisano, F. Lambiase, M. Durante, Polymer self-heating during incremental forming, Journal of Manufacturing Processes. 58 (2020) 1189-1199.
https://doi.org/10.1016/j.jmapro.2020.09.031

[7] L. Li, J. Wang, Equivalent heating tool method for numerical simulation for sheet temperature in frictional stir incremental forming, The International Journal of Advanced Manufacturing Technology. 107 (2020) 1251-1263. https://doi.org/10.1007/s00170-020-05042-1

[8] S. Cai, R. Wu, Z. Wang, M. Li, J. Chen, Numerical simulation of friction stir-assisted incremental forming with synchronous bonding of heterogeneous sheet metals, The International Journal of Advanced Manufacturing Technology. 106 (2020) 2747-2763.
https://doi.org/10.1007/s00170-019-04792-x

[9] N. Duc-Toan, P. Jin-Gee, K. Young-Suk, Combined kinematic/isotropic hardening behavior study for magnesium alloy sheets to predict ductile fracture of rotational incremental forming, International Journal of Material Forming. 3 (2010) 939-942.
https://doi.org/10.1007/s12289-010-0923-9
(2) Although under same processing condition, the surface finish and pattern of springback of the two materials are quite different due to the combined effect of spindle speed and thermal conductivities.

(3) Based on the results, it can be concluded that for material with high conductivity, the FSIF could be a viable option to achieve enhanced formability; whilst for sheet material with low conductivity, it would be better to use a different approach such as global heating (e.g., hot air heating) to obtain similar degree of formability improvement.

## Acknowledgements

This work was supported by the UK Engineering and Physical Science Research Council [Grant No. EP/W010089/1].

## References

[1] H. Zhu, H. Ou, A. Popov, Incremental sheet forming of thermoplastics: a review, The International Journal of Advanced Manufacturing Technology. 111 (2020) 565-587.
https://doi.org/10.1007/s00170-020-06056-5

[2] W. Peng, H. Ou, Deformation mechanisms and fracture in tension under cyclic bending plus compression, single point and double-sided incremental sheet forming processes, International Journal of Machine Tools and Manufacture. 184 (2023) 103980.
https://doi.org/10.1016/j.ijmachtools.2022.103980

[3] M. Li, R. Wu, S. Cai, Z. Chang, Z. Wang, J. Chen, Experimental investigation on friction-stir-assisted incremental forming with synchronous bonding of aluminum alloy and steel sheets, Journal of Materials Engineering and Performance. 29 (2020) 750-759.
https://doi.org/10.1007/s11665-020-04600-8

[4] Z. Wang, S. Cai, J. Chen, Experimental investigations on friction stir assisted single point incremental forming of low-ductility aluminum alloy sheet for higher formability with reasonable surface quality, Journal of Materials Processing Technology. 277 (2020) 116488.
https://doi.org/10.1016/j.jmatprotec.2019.116488

[5] P.A. Grün, E.H. Uheida, L. Lachmann, D. Dimitrov, G.A. Oosthuizen, Formability of titanium alloy sheets by friction stir incremental forming, The International Journal of Advanced Manufacturing Technology. 99 (2018) 1993-2003. https://doi.org/10.1007/s00170-018-2541-5

[6] A. Formisano, F. Lambiase, M. Durante, Polymer self-heating during incremental forming, Journal of Manufacturing Processes. 58 (2020) 1189-1199.
https://doi.org/10.1016/j.jmapro.2020.09.031

[7] L. Li, J. Wang, Equivalent heating tool method for numerical simulation for sheet temperature in frictional stir incremental forming, The International Journal of Advanced Manufacturing Technology. 107 (2020) 1251-1263. https://doi.org/10.1007/s00170-020-05042-1

[8] S. Cai, R. Wu, Z. Wang, M. Li, J. Chen, Numerical simulation of friction stir-assisted incremental forming with synchronous bonding of heterogeneous sheet metals, The International Journal of Advanced Manufacturing Technology. 106 (2020) 2747-2763.
https://doi.org/10.1007/s00170-019-04792-x

[9] N. Duc-Toan, P. Jin-Gee, K. Young-Suk, Combined kinematic/isotropic hardening behavior study for magnesium alloy sheets to predict ductile fracture of rotational incremental forming, International Journal of Material Forming. 3 (2010) 939-942.
https://doi.org/10.1007/s12289-010-0923-9
[10] A. Giuseppina, C. Claudio, F. Luigino, G. Francesco, Theoretical model for temperature prediction in incremental sheet forming-experimental validation, International Journal of Mechanical Sciences. 108 (2016) 39-48. https://doi.org/10.1016/j.ijmecsci.2016.01.030

[11] G. Li, Z. He, J. Ma, H. Yang, H. Li, Springback analysis for warm bending of titanium tube based on coupled thermal-mechanical simulation, Materials. 14 (2021) 5044. https://doi.org/10.3390/ma14175044
[10] A. Giuseppina, C. Claudio, F. Luigino, G. Francesco, Theoretical model for temperature prediction in incremental sheet forming-experimental validation, International Journal of Mechanical Sciences. 108 (2016) 39-48. https://doi.org/10.1016/j.ijmecsci.2016.01.030

[11] G. Li, Z. He, J. Ma, H. Yang, H. Li, Springback analysis for warm bending of titanium tube based on coupled thermal-mechanical simulation, Materials. 14 (2021) 5044. https://doi.org/10.3390/ma14175044
Innovative joining by
forming technologies

<figure><img src="image_1545.png" /></figure>
Innovative joining by
forming technologies

<figure><img src="image_1549.png" /></figure>
# Experimental work on friction riveting process of Ti6Al4V in a CNC machine

TAN Irène¹,ᵃ*, COHEN Guillaume¹,ᵇ, ARAUJO Anna-Carla¹,ᶜ, DAIDIE Alain¹,ᵈ

¹ Institut Clément Ader, Université de Toulouse, CNRS/INSA/ISAE/Mines, Albi/UPS, Toulouse, France

ªirene.tan@insa-toulouse.fr, ᵇguillaume.cohen@iut-tlse3.fr, ᶜaraujo@insa-toulouse.fr,

ᵈalain.daidie@insa-toulouse.fr

**Keywords:** Smart Manufacturing, Monitoring, Metal Forming, Friction Riveting, Ti-6Al-4V

**Abstract.** Friction Forge Riveting (FFR) is now being considered as a good alternative for reinforcing multi-material assemblies in industrial applications with the aim of setting up a robotized assembly process. This paper investigates the feasibility of FFR through experiments using a CNC machine and force measurement instrumentation. Different combinations of parameters such as spindle speed and feed per revolution are tested to understand the forming mechanism of a Ti-6Al-4V rivet by friction. The results show that the evolution of forces in response to the load applied to the rivet by the machine is not constant. An oxide layer is observed on the surface of the rivet for a particular combination of parameters.

## Introduction

Aerospace structures are assembled using a variety of processes. One of them, the riveting process, requires several steps such as drilling, milling, inserting and closing the rivet to achieve the final assembly. Plastic deformation of the rivet head is the most commonly used method to close the assembly by riveting. Special machines, pneumatic or electromagnetic [1], [2], are mainly used. To achieve plastic deformation of the rivet head, a force is applied to the fastener. Its amplitude and trajectory can vary according to the different techniques, but it can cause defects in the assembled sheets and/or in the rivet head, which can lead to fragilization of the assembly. Many studies mention this type of problem.

The impact of the process parameters on the assembly's quality have been studied by Korbel [3] and Cheraghi [4] using numerical simulation. They worked on the influence of the forming force, the rivet's dimensions and the diameter of the hole to show logically that an increase in the riveting force was given with the increase of the rivet's diameter. Xie [5] worked on the effect of ultrasonic vibrations to improve the assembly quality. The vibration can reduce the force needed to form the rivet head. Manes et al [6] studied the stress concentration using Crossland criterion and showed the importance of riveting force to characterize the stress field around the fastener, as this field can impact severely the lifetime of the structure. Lepretre et al [7] worked on a numerical simulation of hot riveting process. Their work aims to improve the stress field around the rivet hole and showed the creation of cracks on the edge beneath the formed rivet head.

An alternative method to conventional riveting process has been proposed by Ni et al [8] to prevent crack during riveting process: Friction Forge Riveting (FFR). This technique presents a great alternative regarding the forces applied during assembly as it is the technique studied in this article. The tool rotates and come in contact over the rivet head, which heats up, and is subjected to plastic deformation as per the tool load at a low feed rate against the rivet. As the rivet heats, this process enables to reduce the forces and consequently the risk of damaging the joint. These principal steps of the FFR are illustrated in Fig. 1: Step 1 and 2 shows that the friction tool turning and moving towards the workpiece, on Step 3 the tool turns in contact to the deformed workpiece without any vertical movement and in Step 4 the tool is removed.

<figure><img src="image_1546.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.
# Experimental work on friction riveting process of Ti6Al4V in a CNC machine

TAN Irène¹,ᵃ*, COHEN Guillaume¹,ᵇ, ARAUJO Anna-Carla¹,ᶜ, DAIDIE Alain¹,ᵈ

¹ Institut Clément Ader, Université de Toulouse, CNRS/INSA/ISAE/Mines, Albi/UPS, Toulouse, France

ªirene.tan@insa-toulouse.fr, ᵇguillaume.cohen@iut-tlse3.fr, ᶜaraujo@insa-toulouse.fr,

ᵈalain.daidie@insa-toulouse.fr

**Keywords:** Smart Manufacturing, Monitoring, Metal Forming, Friction Riveting, Ti-6Al-4V

**Abstract.** Friction Forge Riveting (FFR) is now being considered as a good alternative for reinforcing multi-material assemblies in industrial applications with the aim of setting up a robotized assembly process. This paper investigates the feasibility of FFR through experiments using a CNC machine and force measurement instrumentation. Different combinations of parameters such as spindle speed and feed per revolution are tested to understand the forming mechanism of a Ti-6Al-4V rivet by friction. The results show that the evolution of forces in response to the load applied to the rivet by the machine is not constant. An oxide layer is observed on the surface of the rivet for a particular combination of parameters.

## Introduction

Aerospace structures are assembled using a variety of processes. One of them, the riveting process, requires several steps such as drilling, milling, inserting and closing the rivet to achieve the final assembly. Plastic deformation of the rivet head is the most commonly used method to close the assembly by riveting. Special machines, pneumatic or electromagnetic [1], [2], are mainly used. To achieve plastic deformation of the rivet head, a force is applied to the fastener. Its amplitude and trajectory can vary according to the different techniques, but it can cause defects in the assembled sheets and/or in the rivet head, which can lead to fragilization of the assembly. Many studies mention this type of problem.

The impact of the process parameters on the assembly's quality have been studied by Korbel [3] and Cheraghi [4] using numerical simulation. They worked on the influence of the forming force, the rivet's dimensions and the diameter of the hole to show logically that an increase in the riveting force was given with the increase of the rivet's diameter. Xie [5] worked on the effect of ultrasonic vibrations to improve the assembly quality. The vibration can reduce the force needed to form the rivet head. Manes et al [6] studied the stress concentration using Crossland criterion and showed the importance of riveting force to characterize the stress field around the fastener, as this field can impact severely the lifetime of the structure. Lepretre et al [7] worked on a numerical simulation of hot riveting process. Their work aims to improve the stress field around the rivet hole and showed the creation of cracks on the edge beneath the formed rivet head.

An alternative method to conventional riveting process has been proposed by Ni et al [8] to prevent crack during riveting process: Friction Forge Riveting (FFR). This technique presents a great alternative regarding the forces applied during assembly as it is the technique studied in this article. The tool rotates and come in contact over the rivet head, which heats up, and is subjected to plastic deformation as per the tool load at a low feed rate against the rivet. As the rivet heats, this process enables to reduce the forces and consequently the risk of damaging the joint. These principal steps of the FFR are illustrated in Fig. 1: Step 1 and 2 shows that the friction tool turning and moving towards the workpiece, on Step 3 the tool turns in contact to the deformed workpiece without any vertical movement and in Step 4 the tool is removed.

<figure><img src="image_1550.png" /></figure>
<figure><img src="image_1547.png" /><figcaption>Fig. 1. Principal steps of Friction Forge Riveting process [9]</figcaption></figure>

Ni et al [10] showed the feasibility of the FFR in forming the head of a titanium rivet using a tungsten alloy tool. The tungsten tools are chosen as per their mechanical and temperature limits. Lower forming forces are applied to homogenize the rivet head formation with the least defects possible in the assembly [9]. The influence of the spindle speed was studied closely as it plays an essential role in the head shaping. This brief literature review shows that there are different techniques and numerical approaches to model the riveting process.

The work of Ni et al associating the riveting process is closer to the activities that will be conducted in this study to develop a riveting process that, in the near future, will be implemented in a robot arm.

A previous study done by the same authors of this present article [11] presented the analysis of a first experimental campaign using a complete 2³ factorial design of experiment for two materials (titanium and aluminum), three levels of spindle speed (3000, 3500 and 4000 rpm) and feed rate (1, 2 and 3 mm/min). A numerical simulation has been developed based on the observation of the aluminum rivet forming. This article completes the study, analyzing higher process parameters in titanium to understand the process in higher productivity.

## Materials and methods

The rivets used in the experiments are in Titanium alloy (Ti-6Al-4V), they have 5 mm diameter and length of 20 mm. The choice of this titanium alloy is related to the use in aircraft construction, appreciated for its mechanical performances, corrosion resistance and light weight.

The rivet head is set in contact with a turning tungsten-lanthanum tool. A Tungsten material is used as per their hardness and wear and heat resistance.

The tests were performed using a CNC machine-tool: DMG DMU-85 monoBLOCK. The spindle speed and the vertical displacement are controlled by the machine PLC. The rivet is held by an ER32 collet chuck in a fixture designed for this FFR experiment. A free predrilled steel plate is placed over the collet to protect the surface of the clamper. The workpiece set is fixed on a dynamometer Kistler type 9257B that is clamped on the machine table. A charge amplifier Kistler type 5070A and a National Instruments analogic-digital card are connected to the dynamometer. The signal is recorded with 250 Hz acquisition frequency. The complete experimental set-up is illustrated on Fig. 2.
<figure><img src="image_1551.png" /><figcaption>Fig. 1. Principal steps of Friction Forge Riveting process [9]</figcaption></figure>

Ni et al [10] showed the feasibility of the FFR in forming the head of a titanium rivet using a tungsten alloy tool. The tungsten tools are chosen as per their mechanical and temperature limits. Lower forming forces are applied to homogenize the rivet head formation with the least defects possible in the assembly [9]. The influence of the spindle speed was studied closely as it plays an essential role in the head shaping. This brief literature review shows that there are different techniques and numerical approaches to model the riveting process.

The work of Ni et al associating the riveting process is closer to the activities that will be conducted in this study to develop a riveting process that, in the near future, will be implemented in a robot arm.

A previous study done by the same authors of this present article [11] presented the analysis of a first experimental campaign using a complete 2³ factorial design of experiment for two materials (titanium and aluminum), three levels of spindle speed (3000, 3500 and 4000 rpm) and feed rate (1, 2 and 3 mm/min). A numerical simulation has been developed based on the observation of the aluminum rivet forming. This article completes the study, analyzing higher process parameters in titanium to understand the process in higher productivity.

## Materials and methods

The rivets used in the experiments are in Titanium alloy (Ti-6Al-4V), they have 5 mm diameter and length of 20 mm. The choice of this titanium alloy is related to the use in aircraft construction, appreciated for its mechanical performances, corrosion resistance and light weight.

The rivet head is set in contact with a turning tungsten-lanthanum tool. A Tungsten material is used as per their hardness and wear and heat resistance.

The tests were performed using a CNC machine-tool: DMG DMU-85 monoBLOCK. The spindle speed and the vertical displacement are controlled by the machine PLC. The rivet is held by an ER32 collet chuck in a fixture designed for this FFR experiment. A free predrilled steel plate is placed over the collet to protect the surface of the clamper. The workpiece set is fixed on a dynamometer Kistler type 9257B that is clamped on the machine table. A charge amplifier Kistler type 5070A and a National Instruments analogic-digital card are connected to the dynamometer. The signal is recorded with 250 Hz acquisition frequency. The complete experimental set-up is illustrated on Fig. 2.
<figure><img src="image_1548.png" /><figcaption>Fig. 2. Photos of experimental set up (on the left with a high-speed camera and the dynamometer, and in the right the tool just before touching the rivet – the square protection plate is not attached to the rivet nor the cone)</figcaption></figure>

A high-speed camera is used to capture the images to show the evolution of the head deformation. Photographs taken at various times during the tests are shown in Fig. 3.

<figure><img src="image_1549.png" /><figcaption>Fig. 3. Rivet forming by friction at 3500 rpm and 2 mm/min for a titanium rivet</figcaption></figure>

The spindle speed parameters used in the experiments were clearly higher compared to the previous study (3000, 3500 and 4000 rpm using 1, 2 and 3 mm/min feed rates) which changes considerably the temperatures. The lower spindle speeds will be referred to DoE1.

In this case, the spindle speed values were: 10000 and 15000 rpm. The defined design of experiments includes feed rates 0.5, 2, 4, 6 and 8 mm/min, refered as DoE2. Although, it was not possible to use the lower feed rates (0.5 and 2 mm/min) with 15000 rpm because it was identified a risk of excessive heat generation, and it could not be completed. Table 1 shows the experimental parameters.

<table><caption>Table 1. Design of Experiments</caption><thead><tr><th rowspan="2"></th><th colspan="3">Spindle Speed (rpm)</th><th colspan="5">Feed Rate (mm/min)</th></tr><tr><th>3000</th><th>3500</th><th>4000</th><th>1</th><th>2</th><th>3</th><th>4</th><th>6</th><th>8</th></tr></thead><tbody><tr><th>Low Spindles (DoE1)</th><td>3000</td><td>3500</td><td>4000</td><td>1</td><td>2</td><td>3</td><td>4</td><td>6</td><td>8</td></tr><tr><th>High Spindles (DoE2)</th><td>10000</td><td>15000</td><td>15000</td><td>0.5</td><td>2</td><td>4</td><td>6</td><td>8</td><td>8</td></tr></tbody></table>

## Results and discussion

Fig. 4 shows a typical force and torque evolution during FFR for low (Fig. 4a and 4b) and high spindles speed (Fig. 4c and 4d). The curve profile shows repeated peaks almost periodic with different average time intervals which are not related to the feed rate nor the spindle frequency.

For each experiment, the amplitude and periodicity (pattern) of these curves were analyzed and compared to the behavior of the observed deformation on the high-speed camera to understand the phenomena.
<figure><img src="image_1552.png" /><figcaption>Fig. 2. Photos of experimental set up (on the left with a high-speed camera and the dynamometer, and in the right the tool just before touching the rivet – the square protection plate is not attached to the rivet nor the cone)</figcaption></figure>

A high-speed camera is used to capture the images to show the evolution of the head deformation. Photographs taken at various times during the tests are shown in Fig. 3.

<figure><img src="image_1553.png" /><figcaption>Fig. 3. Rivet forming by friction at 3500 rpm and 2 mm/min for a titanium rivet</figcaption></figure>

The spindle speed parameters used in the experiments were clearly higher compared to the previous study (3000, 3500 and 4000 rpm using 1, 2 and 3 mm/min feed rates) which changes considerably the temperatures. The lower spindle speeds will be referred to DoE1.

In this case, the spindle speed values were: 10000 and 15000 rpm. The defined design of experiments includes feed rates 0.5, 2, 4, 6 and 8 mm/min, refered as DoE2. Although, it was not possible to use the lower feed rates (0.5 and 2 mm/min) with 15000 rpm because it was identified a risk of excessive heat generation, and it could not be completed. Table 1 shows the experimental parameters.

<table><caption>Table 1. Design of Experiments</caption><thead><tr><th rowspan="2"></th><th colspan="3">Spindle Speed (rpm)</th><th colspan="5">Feed Rate (mm/min)</th></tr><tr><th>3000</th><th>3500</th><th>4000</th><th>1</th><th>2</th><th>3</th><th>4</th><th>6</th><th>8</th></tr></thead><tbody><tr><th>Low Spindles (DoE1)</th><td>3000</td><td>3500</td><td>4000</td><td>1</td><td>2</td><td>3</td><td>4</td><td>6</td><td>8</td></tr><tr><th>High Spindles (DoE2)</th><td>10000</td><td>15000</td><td>15000</td><td>0.5</td><td>2</td><td>4</td><td>6</td><td>8</td><td>8</td></tr></tbody></table>

## Results and discussion

Fig. 4 shows a typical force and torque evolution during FFR for low (Fig. 4a and 4b) and high spindles speed (Fig. 4c and 4d). The curve profile shows repeated peaks almost periodic with different average time intervals which are not related to the feed rate nor the spindle frequency.

For each experiment, the amplitude and periodicity (pattern) of these curves were analyzed and compared to the behavior of the observed deformation on the high-speed camera to understand the phenomena.
<figure><img src="image_1550.png" /></figure>

<figure><img src="image_1551.png" /></figure>

<figure><img src="image_1552.png" /></figure>

<figure><img src="image_1553.png" /><figcaption>Fig. 4. Typical evolution of measured data during the friction riveting process at a) 3500 rpm and 2 mm/min, b) 10000 rpm and 2 mm/min</figcaption></figure>

For all experiments, maximum forces and signal periodicity have been compared. Fig. 5 shows maximum forces observed: Fig. 5(a) and (b) show the maximum force and torque for DoE1 and Fig. 5(c) and (d) for DoE2, as a function of the feed rate. It was noted that higher spindle speeds produced lower forces on the rivet with the same tool displacement.

For DoE1, the force evolution seems to mainly increases as the feed rate increases looking at the evolution for parameters 3000 and 3500 rpm but not for 4000 rpm, that remained in the same level of 3500 N. The torque evolution shows the inverse tendency except that this decreasing behavior is observable for parameters 3000 rpm and 4000 rpm, and not for 3500 rpm.

For DoE2, force and torque decrease as the feed rate increases, except at 4 mm/min. At this state, the fluctuation of these variables could be due to a delay in the start of acquisition and so missing data at the beginning as it can be observable in Fig. 4. But this assumption would only be possible for forces measurements.
<figure><img src="image_1554.png" /></figure>

<figure><img src="image_1555.png" /></figure>

<figure><img src="image_1556.png" /></figure>

<figure><img src="image_1557.png" /><figcaption>Fig. 4. Typical evolution of measured data during the friction riveting process at a) 3500 rpm and 2 mm/min, b) 10000 rpm and 2 mm/min</figcaption></figure>

For all experiments, maximum forces and signal periodicity have been compared. Fig. 5 shows maximum forces observed: Fig. 5(a) and (b) show the maximum force and torque for DoE1 and Fig. 5(c) and (d) for DoE2, as a function of the feed rate. It was noted that higher spindle speeds produced lower forces on the rivet with the same tool displacement.

For DoE1, the force evolution seems to mainly increases as the feed rate increases looking at the evolution for parameters 3000 and 3500 rpm but not for 4000 rpm, that remained in the same level of 3500 N. The torque evolution shows the inverse tendency except that this decreasing behavior is observable for parameters 3000 rpm and 4000 rpm, and not for 3500 rpm.

For DoE2, force and torque decrease as the feed rate increases, except at 4 mm/min. At this state, the fluctuation of these variables could be due to a delay in the start of acquisition and so missing data at the beginning as it can be observable in Fig. 4. But this assumption would only be possible for forces measurements.
<figure><img src="image_1554.png" /></figure>

<figure><img src="image_1555.png" /></figure>

<figure><img src="image_1556.png" /></figure>

<figure><img src="image_1557.png" /><figcaption>Fig. 5. Evolution of maximum force compared to feed rate for a) DoE1 and b) DoE2, and evolution of maximum torque for c) DoE1 and d) DoE2</figcaption></figure>

In addition to the maximum forces, the signal patterns were analyzed from each signal. A periodicity of this pattern can be identified, as an average time between each peak. Fig. 6 shows the periodicity for DoE1 and DoE2 as a function of feed rate. Globally, the signal periodicity decreases as the feed rate increases, except for 15000 rpm.

The experiments using 15000 rpm presented a different pattern. These results could be explained by the apparition of an oxide layer which become darker as we increased the feed rate, as illustrated in Fig. 7 comparing the force and torque results for this specific case. Observing Fig. 7, it is as if the oxide apparition has broken the evolution of each variable. This oxide layer could have created a resistance to the forming process, generating the reduction (drop down) of the spindle values. The signal transmission interruption could explain the increase of the signal period. The assumption of a too important heat generated can be mentioned as well.

It would be very important to have temperature acquisition to complete the analysis. Two different set-ups were implemented to measure temperature. Using IR camera, the sparks and the range of measured temperature was not compatible with the different options of the equipment. Using thermocouples on the collet, the data was not trustable and very noisy. So, it was not possible to have trust data during the experiments to verify this assumption.
<figure><img src="image_1558.png" /><figcaption>Fig. 5. Evolution of maximum force compared to feed rate for a) DoE1 and b) DoE2, and evolution of maximum torque for c) DoE1 and d) DoE2</figcaption></figure>

In addition to the maximum forces, the signal patterns were analyzed from each signal. A periodicity of this pattern can be identified, as an average time between each peak. Fig. 6 shows the periodicity for DoE1 and DoE2 as a function of feed rate. Globally, the signal periodicity decreases as the feed rate increases, except for 15000 rpm.

The experiments using 15000 rpm presented a different pattern. These results could be explained by the apparition of an oxide layer which become darker as we increased the feed rate, as illustrated in Fig. 7 comparing the force and torque results for this specific case. Observing Fig. 7, it is as if the oxide apparition has broken the evolution of each variable. This oxide layer could have created a resistance to the forming process, generating the reduction (drop down) of the spindle values. The signal transmission interruption could explain the increase of the signal period. The assumption of a too important heat generated can be mentioned as well.

It would be very important to have temperature acquisition to complete the analysis. Two different set-ups were implemented to measure temperature. Using IR camera, the sparks and the range of measured temperature was not compatible with the different options of the equipment. Using thermocouples on the collet, the data was not trustable and very noisy. So, it was not possible to have trust data during the experiments to verify this assumption.
<figure><img src="image_1558.png" /><figcaption>b) Period [s]</figcaption></figure>

<figure><img src="image_1559.png" /><figcaption>Fig. 6. Evolution of the force signal periodicity compared to feed rate for a) DoE1 and b) DoE2</figcaption></figure>

<figure><img src="image_1560.png" /><figcaption>Fig. 7. Evolution of the force signal periodicity compared to feed rate at 15000 rpm (DoE2)</figcaption></figure>

## Conclusion and perspectives

Friction forming is a complex forming process which implies heat generation and potentially oxidation of the titanium head. Experimental tests of frictional riveting head formation were performed using special parameters that were not experimented before in the literature. The experimental acquisition (forces, torque and high-speed camera) during the tests allowed to observe and understand the rivet forming process. The results highlighted the influence of the spindle speed and feed rate on the behavior of forces on the rivet head using controlled vertical displacement. The force curves presented a special pattern that were not observed in the literature. Semi-periodic results allowed to identify the period of forces and to compare with the physical behavior of the rivet. These results should be analyzed to compare with the external sensor data. Temperature should be measured in a further study to allow numerical simulation studies. Further experiments are being developed, including temperature measurements, power and current (torque) data taken from the CNC PLC, to construct a digital twin with numerical simulation of the titanium forming rivet head.
<figure><img src="image_1559.png" /><figcaption>b) Period [s]</figcaption></figure>

<figure><img src="image_1560.png" /><figcaption>Fig. 6. Evolution of the force signal periodicity compared to feed rate for a) DoE1 and b) DoE2</figcaption></figure>

<figure><img src="image_1561.png" /><figcaption>Fig. 7. Evolution of the force signal periodicity compared to feed rate at 15000 rpm (DoE2)</figcaption></figure>

## Conclusion and perspectives

Friction forming is a complex forming process which implies heat generation and potentially oxidation of the titanium head. Experimental tests of frictional riveting head formation were performed using special parameters that were not experimented before in the literature. The experimental acquisition (forces, torque and high-speed camera) during the tests allowed to observe and understand the rivet forming process. The results highlighted the influence of the spindle speed and feed rate on the behavior of forces on the rivet head using controlled vertical displacement. The force curves presented a special pattern that were not observed in the literature. Semi-periodic results allowed to identify the period of forces and to compare with the physical behavior of the rivet. These results should be analyzed to compare with the external sensor data. Temperature should be measured in a further study to allow numerical simulation studies. Further experiments are being developed, including temperature measurements, power and current (torque) data taken from the CNC PLC, to construct a digital twin with numerical simulation of the titanium forming rivet head.
## Acknowledgement

This work was carried out with the support of the DGA in partnership with the companies Avantis Project and Latécoère.

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[8] R. Ni, W. Hou, Y. Shen, W. Liu, F. Cao, et T. Sun, « Friction forge riveting: A new joining method for connecting 40Cr steel and TC4 titanium alloy », *J. Manuf. Process.*, vol. 68, p. 79-89, juill. 2021. https://doi.org/10.1016/j.jmapro.2021.07.008

[9] R. Ni et al., « Simulation-based parameter optimization of friction forge riveting for AA6061-T6 and TA2 with TA2 titanium rivet », *J. Manuf. Process.*, vol. 83, p. 1-13, août 2022. https://doi.org/10.1016/j.jmapro.2022.08.027

[10] R. Ni, L. Liu, Y. Shen, F. Cao, Y. Yan, et W. Liu, « Friction forge riveting of AA6061-T6 and TA2 plates with large diameter TA2 titanium rivets », *J. Mater. Process. Technol.*, vol. 294, p. 117119, févr. 2021. https://doi.org/10.1016/j.jmatprotec.2021.117119

[11] I. Tan, A. Daidié, G. Cohen, et A.-C. Araujo, « Study of the Friction Forge Riveting (FFR) Process and Numerical Simulation », in *Proceedings of the 14th International Conference on the Technology of Plasticity - Current Trends in the Technology of Plasticity*, K. Mocellin, P.-O. Bouchard, R. Bigot, et T. Balan, Éd., in Lecture Notes in Mechanical Engineering., Cham: Springer Nature Switzerland, 2024, p. 190-200. doi: 10.1007/978-3-031-41341-4_20
## Acknowledgement

This work was carried out with the support of the DGA in partnership with the companies Avantis Project and Latécoère.

## References

[1] E. A. Repetto, R. Radovitzky, M. Ortiz, R. C. Lundquist, et D. R. Sandstrom, « A Finite Element Study of Electromagnetic Riveting », *J. Manuf. Sci. Eng.*, vol. 121, n° 1, p. 61-68, févr. 1999. https://doi.org/10.1115/1.2830576

[2] Z. Cao et Y. Zuo, « Electromagnetic riveting technique and its applications », *Chin. J. Aeronaut.*, vol. 33, n° 1, p. 5-15, janv. 2020. https://doi.org/10.1016/j.cja.2018.12.023

[3] A. Korbel, « Effect of aircraft rivet installation process and production variables on residual stress, clamping force and fatigue behaviour of thin sheet riveted lap joints », *Thin-Walled Struct.*, vol. 181, 2022. https://doi.org/10.1016/j.tws.2022.110041

[4] S. H. Cheraghi, « Effect of variations in the riveting process on the quality of riveted joints », *Int. J. Adv. Manuf. Technol.*, vol. 39, n° 11-12, p. 1144-1155, déc. 2008. https://doi.org/10.1007/s00170-007-1291-6

[5] Z. Xie, F. Chen, et W. He, « The effects of ultrasonic vibration on riveting quality », *Sci. Rep.*, vol. 12, juill. 2022. https://doi.org/10.1038/s41598-022-17095-1

[6] A. Manes, M. Giglio, et F. Viganò, « Effect of riveting process parameters on the local stress field of a T-joint », *Int. J. Mech. Sci.*, vol. 53, n° 12, p. 1039-1049, déc. 2011. https://doi.org/10.1016/j.ijmecsci.2011.07.013

[7] E. Lepretre, S. Chataigner, L. Dieng, et H. Cannard, « Numerical and experimental investigations of hot driven riveting process on old metal structures | Elsevier Enhanced Reader », *Eng. Struct.*, vol. 127, p. 583-593, avr. 2016. https://doi.org/10.1016/j.engstruct.2016.08.056

[8] R. Ni, W. Hou, Y. Shen, W. Liu, F. Cao, et T. Sun, « Friction forge riveting: A new joining method for connecting 40Cr steel and TC4 titanium alloy », *J. Manuf. Process.*, vol. 68, p. 79-89, juill. 2021. https://doi.org/10.1016/j.jmapro.2021.07.008

[9] R. Ni et al., « Simulation-based parameter optimization of friction forge riveting for AA6061-T6 and TA2 with TA2 titanium rivet », *J. Manuf. Process.*, vol. 83, p. 1-13, août 2022. https://doi.org/10.1016/j.jmapro.2022.08.027

[10] R. Ni, L. Liu, Y. Shen, F. Cao, Y. Yan, et W. Liu, « Friction forge riveting of AA6061-T6 and TA2 plates with large diameter TA2 titanium rivets », *J. Mater. Process. Technol.*, vol. 294, p. 117119, févr. 2021. https://doi.org/10.1016/j.jmatprotec.2021.117119

[11] I. Tan, A. Daidié, G. Cohen, et A.-C. Araujo, « Study of the Friction Forge Riveting (FFR) Process and Numerical Simulation », in *Proceedings of the 14th International Conference on the Technology of Plasticity - Current Trends in the Technology of Plasticity*, K. Mocellin, P.-O. Bouchard, R. Bigot, et T. Balan, Éd., in Lecture Notes in Mechanical Engineering., Cham: Springer Nature Switzerland, 2024, p. 190-200. doi: 10.1007/978-3-031-41341-4_20
# Form-based manufacturing of aluminium and steel auxiliary joining elements as the basis for an efficient joining operation

BORGERT Thomas¹,ᵃ*, NORDIEKER Ansgar¹,ᵇ and HOMBERG Werner¹,ᶜ

¹Forming and Machining Technology, Paderborn University, Warburger Straße 100, 33098 Paderborn, Germany

ᵃtb@luf.upb.de, ᵇan@luf.upb.de, ᶜwh@luf.upb.de

**Keywords:** Joining, Thermomechanical, Steel, Aluminium

**Abstract.** Reducing the weight of vehicles can significantly lower the energy or fuel consumed and thus the emissions during operation. One possibility to assess this is the use of a property adapted multi-material systems containing high strength steel, light metals like aluminium or magnesium and fibre reinforced plastics. While expanding the number of materials used new challenges arise for the production and furthermore the joining technology to manufacture the vehicle made of the multi-material systems. One approach to overcome these challenges is to use innovative and adaptable joining techniques which allows the manufacturing of joints of different material combinations. Extensive research activities on the two stage thermo-mechanical joining process with adaptable joining elements was able to demonstrate the great potentials in terms of joining dissimilar materials with good strength. The previously kinematic and path-based fabrication of auxiliary joining elements is modified in this publication to a form-based approach with a perspective of establishing an efficient process chain using easily and cheaply available rods. Based on the new approach to produce the auxiliary joining elements, it can be demonstrated that a reproducible production of the geometry is possible for the investigated steel as well as aluminium material.

**Introduction**

Production technology has to face a wide variety of challenges in the future. One example of this is to reduce the effects and consequences of climate change as far as possible through a massive reduction in CO2 emissions [1,2]. In addition, various crises with worldwide effects lead to significantly increased prices as well as restrictions in the availability of materials and resources or the ability to deliver. All these challenges influence the various sub-areas of production technology, especially joining technology. The challenges described above with respect to modern joining technology result, among other things, in two different requirements that cannot be met by all conventional joining processes [3]. On the one hand, savings in CO2 emissions in the transport sector are being implemented through the consistent lightweight construction of vehicles, which is leading to an increase in the materials used in the vehicle and the resulting material combinations [4,5]. Many of the classic and conventional joining processes are not or only poorly suited for joining these different material combinations. Although thermal joining processes are suitable for joining different metals (e. g. aluminum and steel sheets) despite their different physical properties, it is not always possible to join different groups of materials (e.g. aluminum with fiber composites) [6].

In recent years, thermomechanical joining by means of the friction stir welding process (FSW) has been a frequent subject of investigation [7]. The friction stir welding process is classified in the main group of friction welding processes (FRW), whereby the central feature of friction stir welding is the use of a tool [8]. In friction stir welding, direct frictional contact between the tool and the sheets to be joined generates the heat necessary for the process to produce the joint between the joining partners [9]. One advantage of this method is the fact that a tool but no auxiliary joining

<figure><img src="image_1561.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.
# Form-based manufacturing of aluminium and steel auxiliary joining elements as the basis for an efficient joining operation

BORGERT Thomas¹,ᵃ*, NORDIEKER Ansgar¹,ᵇ and HOMBERG Werner¹,ᶜ

¹Forming and Machining Technology, Paderborn University, Warburger Straße 100, 33098 Paderborn, Germany

ᵃtb@luf.upb.de, ᵇan@luf.upb.de, ᶜwh@luf.upb.de

**Keywords:** Joining, Thermomechanical, Steel, Aluminium

**Abstract.** Reducing the weight of vehicles can significantly lower the energy or fuel consumed and thus the emissions during operation. One possibility to assess this is the use of a property adapted multi-material systems containing high strength steel, light metals like aluminium or magnesium and fibre reinforced plastics. While expanding the number of materials used new challenges arise for the production and furthermore the joining technology to manufacture the vehicle made of the multi-material systems. One approach to overcome these challenges is to use innovative and adaptable joining techniques which allows the manufacturing of joints of different material combinations. Extensive research activities on the two stage thermo-mechanical joining process with adaptable joining elements was able to demonstrate the great potentials in terms of joining dissimilar materials with good strength. The previously kinematic and path-based fabrication of auxiliary joining elements is modified in this publication to a form-based approach with a perspective of establishing an efficient process chain using easily and cheaply available rods. Based on the new approach to produce the auxiliary joining elements, it can be demonstrated that a reproducible production of the geometry is possible for the investigated steel as well as aluminium material.

**Introduction**

Production technology has to face a wide variety of challenges in the future. One example of this is to reduce the effects and consequences of climate change as far as possible through a massive reduction in CO2 emissions [1,2]. In addition, various crises with worldwide effects lead to significantly increased prices as well as restrictions in the availability of materials and resources or the ability to deliver. All these challenges influence the various sub-areas of production technology, especially joining technology. The challenges described above with respect to modern joining technology result, among other things, in two different requirements that cannot be met by all conventional joining processes [3]. On the one hand, savings in CO2 emissions in the transport sector are being implemented through the consistent lightweight construction of vehicles, which is leading to an increase in the materials used in the vehicle and the resulting material combinations [4,5]. Many of the classic and conventional joining processes are not or only poorly suited for joining these different material combinations. Although thermal joining processes are suitable for joining different metals (e. g. aluminum and steel sheets) despite their different physical properties, it is not always possible to join different groups of materials (e.g. aluminum with fiber composites) [6].

In recent years, thermomechanical joining by means of the friction stir welding process (FSW) has been a frequent subject of investigation [7]. The friction stir welding process is classified in the main group of friction welding processes (FRW), whereby the central feature of friction stir welding is the use of a tool [8]. In friction stir welding, direct frictional contact between the tool and the sheets to be joined generates the heat necessary for the process to produce the joint between the joining partners [9]. One advantage of this method is the fact that a tool but no auxiliary joining

<figure><img src="image_1562.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.
elements are required for the process [8]. A disadvantage is that the joint strength produced is strongly dependent on the process parameters used as well as the properties of the joining partners used (such as thermal conductivity), so that extensive process control is necessary [10]. Another remedy to the challenge of joining multi-material joints that occurs in thermal joining processes can be the use of mechanical joining processes, which, however, also have different requirements for the materials to be processed depending on the process (such as sufficient ductility of the sheets in clinching) [11,12].

A second challenge for joining technology is posed by fluctuating sheet thicknesses and properties (e.g., due to problems and challenges in supply chains caused by crises) as well as the increasing variety of industrial products. This means that the joining process must become more versatile to be able to react to the changing properties, even between several joining operations [13].

One group of joining processes (which is also well suited to the joining of dissimilar material pairs) are the mechanical joining processes [14]. In addition to improving the versatility of conventional and already extensively researched joining processes such as semi-tubular self-piercing riveting or clinching, extensive research is being carried out in thermomechanical forming processes, which are characterised by the combination of forming and the application of heat to reduce the acting forces and tool loads. The general feasibility of the process to produce joints with a wide range of material combinations and sheet thickness combinations with Friction Spun Joint Connectors (FSJC) as well as good joint strengths has already been successfully demonstrated in past investigations. The entire process sequence can be described in two sub-processes, the thermo-mechanical forming of the Friction Spun Joint Connectors made of aluminium and steel materials as the first step and the production of the joint using the previously produced FJSC as the second step [15]. Despite the extensive work in the past, there is still potential for optimisation for further improvement and development of an industrial application, especially in the first step of the FSJC production.

The process sequence used to this point and the associated tool components used to manufacture the FSJC are shown in Figure 1. Here the tool consists of a sintered carbide KX 40 (9% Cobalt, 91% tungsten carbide) to simultaneously have high mechanical strength and temperature resistance at the same time. The basis of the first process step is the clamping of a round bar section (e. g. made of aluminium or steel materials) in the collet of a high-speed spindle with a maximum rotational speed of n = 15,000 rpm as well as the positioning of the carbide tool. Subsequently This is followed by a rubbing phase, the aim of which is to achieve friction between the tool and the round bar section to heat the latter, reduce the strength to reduce the mechanical loads on the tool. The actual forming of the desired FSJC geometry represents the third and therefore final process step. One great advantage of this kinematic of the FSJC is the ability to specifically influence and adapt the geometry (both the shank length and the diameter) by changing the trajectory of the forming die. This enables the targeted FSJC geometry between two different joining points, which can be caused for example due to fluctuations in production or a high variance in the joints.

Regardless of this high level of versatility (compared to many other mechanical joining processes), past investigations have also revealed the challenges in this thermomechanical forming process. These challenges are primarily found in the processing of steel materials (e. g. C45E), the processing of aluminium materials is without problems possible [16].

When processing the heat-treatable steel C45E, abrasive wear of the tools occurs due to the high process temperatures (above 1000 °C), which leads to inconsistent FSJC geometries. In addition, the setting of a flange inclined on the sheet side (which is necessary to realise a force fit) is only possible for defined parameter combinations. Another challenge to be solved for the industrial implementation of the process is the process kinematics. The process sequence described above (and shown in Figure 1) requires the separation and individual clamping of the semi-finished
elements are required for the process [8]. A disadvantage is that the joint strength produced is strongly dependent on the process parameters used as well as the properties of the joining partners used (such as thermal conductivity), so that extensive process control is necessary [10]. Another remedy to the challenge of joining multi-material joints that occurs in thermal joining processes can be the use of mechanical joining processes, which, however, also have different requirements for the materials to be processed depending on the process (such as sufficient ductility of the sheets in clinching) [11,12].

A second challenge for joining technology is posed by fluctuating sheet thicknesses and properties (e.g., due to problems and challenges in supply chains caused by crises) as well as the increasing variety of industrial products. This means that the joining process must become more versatile to be able to react to the changing properties, even between several joining operations [13].

One group of joining processes (which is also well suited to the joining of dissimilar material pairs) are the mechanical joining processes [14]. In addition to improving the versatility of conventional and already extensively researched joining processes such as semi-tubular self-piercing riveting or clinching, extensive research is being carried out in thermomechanical forming processes, which are characterised by the combination of forming and the application of heat to reduce the acting forces and tool loads. The general feasibility of the process to produce joints with a wide range of material combinations and sheet thickness combinations with Friction Spun Joint Connectors (FSJC) as well as good joint strengths has already been successfully demonstrated in past investigations. The entire process sequence can be described in two sub-processes, the thermo-mechanical forming of the Friction Spun Joint Connectors made of aluminium and steel materials as the first step and the production of the joint using the previously produced FJSC as the second step [15]. Despite the extensive work in the past, there is still potential for optimisation for further improvement and development of an industrial application, especially in the first step of the FSJC production.

The process sequence used to this point and the associated tool components used to manufacture the FSJC are shown in Figure 1. Here the tool consists of a sintered carbide KX 40 (9% Cobalt, 91% tungsten carbide) to simultaneously have high mechanical strength and temperature resistance at the same time. The basis of the first process step is the clamping of a round bar section (e. g. made of aluminium or steel materials) in the collet of a high-speed spindle with a maximum rotational speed of n = 15,000 rpm as well as the positioning of the carbide tool. Subsequently This is followed by a rubbing phase, the aim of which is to achieve friction between the tool and the round bar section to heat the latter, reduce the strength to reduce the mechanical loads on the tool. The actual forming of the desired FSJC geometry represents the third and therefore final process step. One great advantage of this kinematic of the FSJC is the ability to specifically influence and adapt the geometry (both the shank length and the diameter) by changing the trajectory of the forming die. This enables the targeted FSJC geometry between two different joining points, which can be caused for example due to fluctuations in production or a high variance in the joints.

Regardless of this high level of versatility (compared to many other mechanical joining processes), past investigations have also revealed the challenges in this thermomechanical forming process. These challenges are primarily found in the processing of steel materials (e. g. C45E), the processing of aluminium materials is without problems possible [16].

When processing the heat-treatable steel C45E, abrasive wear of the tools occurs due to the high process temperatures (above 1000 °C), which leads to inconsistent FSJC geometries. In addition, the setting of a flange inclined on the sheet side (which is necessary to realise a force fit) is only possible for defined parameter combinations. Another challenge to be solved for the industrial implementation of the process is the process kinematics. The process sequence described above (and shown in Figure 1) requires the separation and individual clamping of the semi-finished
product sections. To improve process and cost efficiency, a long wire segment must be used and continuously fed. This requires a reversal of the kinematics in relation to the rotating and stationary process elements.

To summarise the challenges, an alternative process strategy must therefore be developed and investigated as part of the investigations described in this publication in order to improve the shaping of the FJSC, reduce tool wear and implement a reversal of the kinematics to avoid cutting individual round bar sections. For this purpose, the approach used with a shaping tool system is investigated for the processing of both aluminium and steel round bars and the results of test series are considered.

<figure><img src="image_1562.png" /><figcaption>Figure 1: Based on the kinematic shaping of a carbide tool, the (a) schematic process principle and (b) real test setup used.</figcaption></figure>

## Materials and Methods

The investigations described in this paper use a modified test setup (compared to the one described in the introduction) (see Figure 2) with modified kinematics. Here, the tool is located in the collet of the horizontally arranged milling spindle (WEISS Spindeltechnologie GmbH, Maroldsweisach, Germany) and is set in rotation during the process. Although the round bar section is also clamped in a collet, it is not rotated and is only fed in the positive y-direction by a cross support. For this modified process kinematics, a rubbing phase is also used to inject heat into the round bar section to reduce the strength and therefore also the tool loads during the thermomechanical forming process. The actual forming process takes place by upsetting the round bar section on the forming tool, in which the shank of the FSJC is produced by the hole and the flange by the flank in the forming tool.

<figure><img src="image_1563.png" /><figcaption>Figure 2: Use of a shaping tool system with (a) schematic process principle and (b) real test setup.</figcaption></figure>
product sections. To improve process and cost efficiency, a long wire segment must be used and continuously fed. This requires a reversal of the kinematics in relation to the rotating and stationary process elements.

To summarise the challenges, an alternative process strategy must therefore be developed and investigated as part of the investigations described in this publication in order to improve the shaping of the FJSC, reduce tool wear and implement a reversal of the kinematics to avoid cutting individual round bar sections. For this purpose, the approach used with a shaping tool system is investigated for the processing of both aluminium and steel round bars and the results of test series are considered.

<figure><img src="image_1563.png" /><figcaption>Figure 1: Based on the kinematic shaping of a carbide tool, the (a) schematic process principle and (b) real test setup used.</figcaption></figure>

## Materials and Methods

The investigations described in this paper use a modified test setup (compared to the one described in the introduction) (see Figure 2) with modified kinematics. Here, the tool is located in the collet of the horizontally arranged milling spindle (WEISS Spindeltechnologie GmbH, Maroldsweisach, Germany) and is set in rotation during the process. Although the round bar section is also clamped in a collet, it is not rotated and is only fed in the positive y-direction by a cross support. For this modified process kinematics, a rubbing phase is also used to inject heat into the round bar section to reduce the strength and therefore also the tool loads during the thermomechanical forming process. The actual forming process takes place by upsetting the round bar section on the forming tool, in which the shank of the FSJC is produced by the hole and the flange by the flank in the forming tool.

<figure><img src="image_1564.png" /><figcaption>Figure 2: Use of a shaping tool system with (a) schematic process principle and (b) real test setup.</figcaption></figure>
To validate the suitability and demonstrate the optimisation potential of the shaping tool system, both round bars (d = 8 mm) of the wrought aluminium alloy EN AW-6060 and the steel alloy C45E are examined in an independent series of tests. The shaping tool here also consists of the sintered carbide KX 40. In the respective test series (limited by the tool geometry), the length of the shank of the FSJC is varied in three stages for both steel and aluminium (l₁ = 6 mm, l₂ = 9 mm, l₃ = 12 mm). A constant shaft diameter of d = 5.2 mm is kept for the steel and aluminum material due to the shaping die.

Preliminary investigations in the processing of the steel material C45E show that the rotational speed has no significant influence on the process, so that only the feed rate during forming is varied in the four stages shown in Table 1.

<table><caption>Table 1: Parameters and their stages for the production of steel FSJC</caption><thead><tr><th colspan="5">Steel</th></tr><tr><th>Stages</th><th>1</th><th>2</th><th>3</th><th>4</th></tr></thead><tbody><tr><th>Parameter</th><td></td><td></td><td></td><td></td></tr><tr><th>Feedrate [mm/min]</th><td>20</td><td>53,3</td><td>86,7</td><td>120</td></tr><tr><th>Shank length [mm]</th><td>6</td><td>9</td><td>12</td><td>/</td></tr></tbody></table>

For the aluminum alloy EN AW-6060, the rotational speed according to the preliminary tests shows a (small) influence on the process and the geometry of the auxiliary joining elements produced. For this reason, the experimental plan varies it over five stages. Due to the lower strength of aluminum compared to steel, higher feed rates can be achieved according to the preliminary tests. Five different feed rates are investigated for the aluminum auxiliary joining elements. The stages for the rotational speed and for the feed rate are shown in Table 2. Flowdrill FDKS release agent paste is used to reduce thermal and mechanical tool stresses during the forming process.

<table><caption>Table 2: Parameters and their stages for the production of aluminium FSJC</caption><thead><tr><th colspan="6">Aluminium</th></tr><tr><th>Stages</th><th>1</th><th>2</th><th>3</th><th>4</th><th>5</th></tr></thead><tbody><tr><th>Parameter</th><td></td><td></td><td></td><td></td><td></td></tr><tr><th>Rotational speed [rpm]</th><td>3000</td><td>3875</td><td>4750</td><td>5625</td><td>6500</td></tr><tr><th>Feedrate [mm/min]</th><td>20</td><td>77,5</td><td>135</td><td>195,5</td><td>250</td></tr><tr><th>Shank length [mm]</th><td>6</td><td>9</td><td>12</td><td>/</td><td>/</td></tr></tbody></table>

The produced FSJCs are extensively characterised after production and the results are compared with those of the kinematic shaping approach. For this purpose, both the weight of the semi-finished product used before the forming process and the weight of the formed FSJC are measured. In addition (based on Figure 3), the actual geometry of the formed FSJC is recorded using tactile measuring devices (calipers and micrometers) and compared with the target geometry.
To validate the suitability and demonstrate the optimisation potential of the shaping tool system, both round bars (d = 8 mm) of the wrought aluminium alloy EN AW-6060 and the steel alloy C45E are examined in an independent series of tests. The shaping tool here also consists of the sintered carbide KX 40. In the respective test series (limited by the tool geometry), the length of the shank of the FSJC is varied in three stages for both steel and aluminium (l₁ = 6 mm, l₂ = 9 mm, l₃ = 12 mm). A constant shaft diameter of d = 5.2 mm is kept for the steel and aluminum material due to the shaping die.

Preliminary investigations in the processing of the steel material C45E show that the rotational speed has no significant influence on the process, so that only the feed rate during forming is varied in the four stages shown in Table 1.

<table><caption>Table 1: Parameters and their stages for the production of steel FSJC</caption><thead><tr><th colspan="5">Steel</th></tr><tr><th>Stages</th><th>1</th><th>2</th><th>3</th><th>4</th></tr></thead><tbody><tr><th>Parameter</th><td></td><td></td><td></td><td></td></tr><tr><th>Feedrate [mm/min]</th><td>20</td><td>53,3</td><td>86,7</td><td>120</td></tr><tr><th>Shank length [mm]</th><td>6</td><td>9</td><td>12</td><td>/</td></tr></tbody></table>

For the aluminum alloy EN AW-6060, the rotational speed according to the preliminary tests shows a (small) influence on the process and the geometry of the auxiliary joining elements produced. For this reason, the experimental plan varies it over five stages. Due to the lower strength of aluminum compared to steel, higher feed rates can be achieved according to the preliminary tests. Five different feed rates are investigated for the aluminum auxiliary joining elements. The stages for the rotational speed and for the feed rate are shown in Table 2. Flowdrill FDKS release agent paste is used to reduce thermal and mechanical tool stresses during the forming process.

<table><caption>Table 2: Parameters and their stages for the production of aluminium FSJC</caption><thead><tr><th colspan="6">Aluminium</th></tr><tr><th>Stages</th><th>1</th><th>2</th><th>3</th><th>4</th><th>5</th></tr></thead><tbody><tr><th>Parameter</th><td></td><td></td><td></td><td></td><td></td></tr><tr><th>Rotational speed [rpm]</th><td>3000</td><td>3875</td><td>4750</td><td>5625</td><td>6500</td></tr><tr><th>Feedrate [mm/min]</th><td>20</td><td>77,5</td><td>135</td><td>195,5</td><td>250</td></tr><tr><th>Shank length [mm]</th><td>6</td><td>9</td><td>12</td><td>/</td><td>/</td></tr></tbody></table>

The produced FSJCs are extensively characterised after production and the results are compared with those of the kinematic shaping approach. For this purpose, both the weight of the semi-finished product used before the forming process and the weight of the formed FSJC are measured. In addition (based on Figure 3), the actual geometry of the formed FSJC is recorded using tactile measuring devices (calipers and micrometers) and compared with the target geometry.
<figure><img src="image_1564.png" /><figcaption>Figure 3: Geometric characterisation of the blank and the Friction Spun Joint Connector (FJSC) produced.</figcaption></figure>

## Results and Discussion

The central result of the investigations is the fact that the manufacturing of FSJC from both steel and aluminium is easily possible with the previously described shaping tool system and the rotating tool. Phenomena occurring here as well as differences to the conventional die system, in which the FJSC shape is created by the tool path, are described and explained below.

The manufacturing of FSJCs with different geometries (in the case of the tool system used for moulding, with different lengths) is an important prerequisite for the industrial applicability of the process, which is based on a high degree of versatility. The targeted adjustment of different shank lengths has been demonstrated in the investigations presented here (see Figure 4, a)). Three exemplary FJSCs made of the steel material C45E with the three investigated shank lengths of $l_1 = 6$ mm, $l_1 = 9$ mm and $l_1 = 12$ mm are shown there. It is also possible to produce the different shank lengths (as well as intermediate stages) from the aluminium material. A direct comparison between the target and actual geometry of the FSJC shows that regarding the shank diameter (target value $d_1 = 5.2$ mm), there are only deviations in the range of a few hundredths of a millimetre for both the aluminium and the steel HFE. These deviations (which are smaller compared to the conventional manufacturing approach based on the tool's path geometry) can be explained by the shape of the innovative tool. The shank geometry is predetermined by the hole within the tool, which leads to small deviations (which are acceptable in terms of manufacturing tolerances). A regression analysis using Matlab software (The Mathworks) shows that for steel processing, neither the feed rate nor the rivet length has a statistically significant trend or influence (p-value < 5 %) on the deviation of the shank diameter. For aluminium processing, only the shank length has a slight positive influence, so that, for example, an increase in shank thickness of up to 0.04 mm can be generated with increasing shank length.

The evaluation of the results of the shank length ($l_2$) show that longer FSJCs are manufactured for the steel material C45E than required. Only the rivet length itself (and not the process parameter of the feed) has a statistically significant effect here, so that a larger deviation from the target geometry (up to $\Delta l_2 = 1$ mm) is also set for a longer rivet length, which is also in line with expectations. For the aluminium alloy EN AW-6060, the use of a high feed rate leads to a reduction in the shank length (maximum effect for the highest feed rate of $f_5 = 250$ mm/min up to $\Delta l_2 = 0.8$ mm). This corresponds to the expectations that a high feed rate leads to a shorter process time, a shorter friction path and thus also lower thermal stress [16]. Because less temperature is induced due to friction, the aluminium does not heat up as much and is less ductile.
<figure><img src="image_1565.png" /><figcaption>Figure 3: Geometric characterisation of the blank and the Friction Spun Joint Connector (FJSC) produced.</figcaption></figure>

## Results and Discussion

The central result of the investigations is the fact that the manufacturing of FSJC from both steel and aluminium is easily possible with the previously described shaping tool system and the rotating tool. Phenomena occurring here as well as differences to the conventional die system, in which the FJSC shape is created by the tool path, are described and explained below.

The manufacturing of FSJCs with different geometries (in the case of the tool system used for moulding, with different lengths) is an important prerequisite for the industrial applicability of the process, which is based on a high degree of versatility. The targeted adjustment of different shank lengths has been demonstrated in the investigations presented here (see Figure 4, a)). Three exemplary FJSCs made of the steel material C45E with the three investigated shank lengths of $l_1 = 6$ mm, $l_1 = 9$ mm and $l_1 = 12$ mm are shown there. It is also possible to produce the different shank lengths (as well as intermediate stages) from the aluminium material. A direct comparison between the target and actual geometry of the FSJC shows that regarding the shank diameter (target value $d_1 = 5.2$ mm), there are only deviations in the range of a few hundredths of a millimetre for both the aluminium and the steel HFE. These deviations (which are smaller compared to the conventional manufacturing approach based on the tool's path geometry) can be explained by the shape of the innovative tool. The shank geometry is predetermined by the hole within the tool, which leads to small deviations (which are acceptable in terms of manufacturing tolerances). A regression analysis using Matlab software (The Mathworks) shows that for steel processing, neither the feed rate nor the rivet length has a statistically significant trend or influence (p-value < 5 %) on the deviation of the shank diameter. For aluminium processing, only the shank length has a slight positive influence, so that, for example, an increase in shank thickness of up to 0.04 mm can be generated with increasing shank length.

The evaluation of the results of the shank length ($l_2$) show that longer FSJCs are manufactured for the steel material C45E than required. Only the rivet length itself (and not the process parameter of the feed) has a statistically significant effect here, so that a larger deviation from the target geometry (up to $\Delta l_2 = 1$ mm) is also set for a longer rivet length, which is also in line with expectations. For the aluminium alloy EN AW-6060, the use of a high feed rate leads to a reduction in the shank length (maximum effect for the highest feed rate of $f_5 = 250$ mm/min up to $\Delta l_2 = 0.8$ mm). This corresponds to the expectations that a high feed rate leads to a shorter process time, a shorter friction path and thus also lower thermal stress [16]. Because less temperature is induced due to friction, the aluminium does not heat up as much and is less ductile.
<figure><img src="image_1565.png" /><figcaption>Figure 4: Produced steel joining elements in a) different lengths and b) in a comparison between the form-based manufacturing (left) and the conventional process (right)</figcaption></figure>

In addition, the auxiliary joining elements produced with the shaping-tool system show a lower weight reduction. The difference in the weight of the Rod before the process compared to the manufactured auxiliary joining element after the process is only a few hundredths of a gram. The small deviations are due to inaccuracies in the scale used. In comparison, up to 40% weight reduction can be expected with conventional auxiliary joining element production. This means that volume constancy can be assumed for the shaping-tool process.

## Summary

The results presented in this paper show an innovative approach to the production of adaptive auxiliary joining elements. These can be used in a multi-material system to solve complex joining problems. The advantage of the auxiliary joining elements lies in the geometry adjustment options. Shank length can be specifically controlled via the process parameters. The diameter is set adaptively via a tool change. This allows a wide variety of sheet thicknesses to be joined. The results have shown that both steel and aluminium auxiliary joining elements can be produced reproducibly. The advantage of the shape-based approach is that tool wear decreases during the process because the loads during the forming process are distributed on an area and not on a single point. In addition, the investigations have shown that the deviation of the shank length of steel auxiliary joining elements only depends on the desired shank length. Due to the volume constancy of the process, the shank increases in length during forming because not all of the material volume can be displaced into the flange. The diameter of the manufactured auxiliary joining elements only depends on the tool due to the innovative approach and is acceptable in relation to manufacturing tolerances.

In future, the kinematics of the joining process will also be changed so that the die rotates and the auxiliary joining element remains at rest. In combination with the investigations from this paper, an industrial process design for manufacturing and joining with adaptive auxiliary joining elements should be developed.

## Acknowledgments

Funded by the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation) – TRR 285/2 – Project-ID 418701707, subproject C03.

## References

[1] W. Cai, K.-H. Lai, C. Liu, F. Wei, M. Ma, S. Jia, Z. Jiang; L. Lv, Promoting sustainability of manufacturing industry through the lean energy-saving and emission-reduction strategy, Sci. Total Environ. 665 (2019) 23–32. https://doi.org/10.1016/j.scitotenv.2019.02.069

[2] G. Pina Cipriano, A. Ahiya, J.F. dos Santos, P. Vilaça, S.T. Amancio-Filho, Single-phase friction riveting: metallic rivet deformation, temperature evolution, and joint mechanical performance, Weld World 64 (2020) 47–58. https://doi.org/10.1007/s40194-019-00803-3
<figure><img src="image_1566.png" /><figcaption>Figure 4: Produced steel joining elements in a) different lengths and b) in a comparison between the form-based manufacturing (left) and the conventional process (right)</figcaption></figure>

In addition, the auxiliary joining elements produced with the shaping-tool system show a lower weight reduction. The difference in the weight of the Rod before the process compared to the manufactured auxiliary joining element after the process is only a few hundredths of a gram. The small deviations are due to inaccuracies in the scale used. In comparison, up to 40% weight reduction can be expected with conventional auxiliary joining element production. This means that volume constancy can be assumed for the shaping-tool process.

## Summary

The results presented in this paper show an innovative approach to the production of adaptive auxiliary joining elements. These can be used in a multi-material system to solve complex joining problems. The advantage of the auxiliary joining elements lies in the geometry adjustment options. Shank length can be specifically controlled via the process parameters. The diameter is set adaptively via a tool change. This allows a wide variety of sheet thicknesses to be joined. The results have shown that both steel and aluminium auxiliary joining elements can be produced reproducibly. The advantage of the shape-based approach is that tool wear decreases during the process because the loads during the forming process are distributed on an area and not on a single point. In addition, the investigations have shown that the deviation of the shank length of steel auxiliary joining elements only depends on the desired shank length. Due to the volume constancy of the process, the shank increases in length during forming because not all of the material volume can be displaced into the flange. The diameter of the manufactured auxiliary joining elements only depends on the tool due to the innovative approach and is acceptable in relation to manufacturing tolerances.

In future, the kinematics of the joining process will also be changed so that the die rotates and the auxiliary joining element remains at rest. In combination with the investigations from this paper, an industrial process design for manufacturing and joining with adaptive auxiliary joining elements should be developed.

## Acknowledgments

Funded by the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation) – TRR 285/2 – Project-ID 418701707, subproject C03.

## References

[1] W. Cai, K.-H. Lai, C. Liu, F. Wei, M. Ma, S. Jia, Z. Jiang; L. Lv, Promoting sustainability of manufacturing industry through the lean energy-saving and emission-reduction strategy, Sci. Total Environ. 665 (2019) 23–32. https://doi.org/10.1016/j.scitotenv.2019.02.069

[2] G. Pina Cipriano, A. Ahiya, J.F. dos Santos, P. Vilaça, S.T. Amancio-Filho, Single-phase friction riveting: metallic rivet deformation, temperature evolution, and joint mechanical performance, Weld World 64 (2020) 47–58. https://doi.org/10.1007/s40194-019-00803-3
[3] S.A. Choudry, J. Kaspar, U. Alber, D. Landgrebe, Integration of an Assessment Methodology for the Selection of Joining Technologies in Lightweight Engineering, Procedia CIRP 70 (2018) 217-222. https://doi.org/10.1016/j.procir.2018.02.034

[4] M. Haghshenas, A.P. Gerlich, Joining of automotive sheet materials by friction-based welding methods: A review, Engineering Science and Technology, an International Journal 21 (2018) 130-148. https://doi.org/10.1016/j.jestch.2018.02.008

[5] S. Rosenthal, F. Maaß, M. Kamaliev, M. Hahn, S. Gies, A.E. Tekkaya, Lightweight in Automotive Components by Forming Technology, Automot. Innov. 3 (2020) 195-209. https://doi.org/10.1007/s42154-020-00103-3

[6] G. Meschut, M. Merklein, A. Brosius, D. Drummer, L. Fratini, U. Füssel, M. Gude, W. Homberg, P. Martins, M. Bobbert, et al., Review on mechanical joining by plastic deformation, Journal of Advanced Joining Processes 5 (2022) 100113. https://doi.org/10.1016/j.jajp.2022.100113

[7] G. Çam, S. Mistikoglu, Recent Developments in Friction Stir Welding of Al-alloys, J. of Materi. Eng. and Perform. 23 (2014) 1936-1953. https://doi.org/10.1007/s11665-014-0968-x

[8] D. Kumar Rajak, D.D. Pagar, P.L. Menezes, A. Eyvazian, Friction-based welding processes: friction welding and friction stir welding, Journal of Adhesion Science and Technology 34 (2020) 2613-2637. https://doi.org/10.1080/01694243.2020.1780716

[9] A. Heidarzadeh, S. Mironov, R. Kaibyshev, G. Çam, A. Simar, A. Gerlich, F. Khodabakhshi, A. Mostafaei, D.P. Field, J.D. Robson, et al., Friction stir welding/processing of metals and alloys: A comprehensive review on microstructural evolution, Progress in Materials Science 117 (2021) 100752. https://doi.org/10.1016/j.pmatsci.2020.100752

[10] A. Heidarzadeh, Tensile behavior, microstructure, and substructure of the friction stir welded 70/30 brass joints: RSM, EBSD, and TEM study, Archives of Civil and Mechanical Engineering 19 (2019) 137-146. https://doi.org/10.1016/j.acme.2018.09.009

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

[12] F. Lambiase, S.I. Scipioni, C.-J. Lee, D.-C. Ko, F. Liu, A State-of-the-Art Review on Advanced Joining Processes for Metal-Composite and Metal-Polymer Hybrid Structures, Materials (Basel) 14 (2021). https://doi.org/10.3390/ma14081890

[13] M. Merklein, M. Jäckisch, C.-M. Kuball, D. Römisch, S. Wiesenmayer, S. Wituschek, Mechanical joining of high-strength multi-material systems - trends and innovations, Mechanics & Industry 24 (2023). https://doi.org/10.1051/meca/2023013

[14] C. Wischer, W. Homberg, A contribution on versatile process chains: joining with adaptive joining elements, formed by friction spinning, Prod. Eng. Res. Devel. 16 (2022) 379-388. https://doi.org/10.1007/s11740-021-01094-8

[15] C. Wischer, W. Homberg, Further Development of an Adaptive Joining Technique Based on Friction Spinning to Produce Pre-Hole-Free Joints, KEM 926 (2022) 1468-1478. https://doi.org/10.4028/p-1n6741

[16] T. Borgert, M. Henke, W. Homberg, Investigations on the Influences of the Thermomechanical Manufacturing of Aluminium Auxiliary Joining Elements, JMMP 7 (2023) 147. https://doi.org/10.3390/jmmp7040147
[3] S.A. Choudry, J. Kaspar, U. Alber, D. Landgrebe, Integration of an Assessment Methodology for the Selection of Joining Technologies in Lightweight Engineering, Procedia CIRP 70 (2018) 217-222. https://doi.org/10.1016/j.procir.2018.02.034

[4] M. Haghshenas, A.P. Gerlich, Joining of automotive sheet materials by friction-based welding methods: A review, Engineering Science and Technology, an International Journal 21 (2018) 130-148. https://doi.org/10.1016/j.jestch.2018.02.008

[5] S. Rosenthal, F. Maaß, M. Kamaliev, M. Hahn, S. Gies, A.E. Tekkaya, Lightweight in Automotive Components by Forming Technology, Automot. Innov. 3 (2020) 195-209. https://doi.org/10.1007/s42154-020-00103-3

[6] G. Meschut, M. Merklein, A. Brosius, D. Drummer, L. Fratini, U. Füssel, M. Gude, W. Homberg, P. Martins, M. Bobbert, et al., Review on mechanical joining by plastic deformation, Journal of Advanced Joining Processes 5 (2022) 100113. https://doi.org/10.1016/j.jajp.2022.100113

[7] G. Çam, S. Mistikoglu, Recent Developments in Friction Stir Welding of Al-alloys, J. of Materi. Eng. and Perform. 23 (2014) 1936-1953. https://doi.org/10.1007/s11665-014-0968-x

[8] D. Kumar Rajak, D.D. Pagar, P.L. Menezes, A. Eyvazian, Friction-based welding processes: friction welding and friction stir welding, Journal of Adhesion Science and Technology 34 (2020) 2613-2637. https://doi.org/10.1080/01694243.2020.1780716

[9] A. Heidarzadeh, S. Mironov, R. Kaibyshev, G. Çam, A. Simar, A. Gerlich, F. Khodabakhshi, A. Mostafaei, D.P. Field, J.D. Robson, et al., Friction stir welding/processing of metals and alloys: A comprehensive review on microstructural evolution, Progress in Materials Science 117 (2021) 100752. https://doi.org/10.1016/j.pmatsci.2020.100752

[10] A. Heidarzadeh, Tensile behavior, microstructure, and substructure of the friction stir welded 70/30 brass joints: RSM, EBSD, and TEM study, Archives of Civil and Mechanical Engineering 19 (2019) 137-146. https://doi.org/10.1016/j.acme.2018.09.009

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

[12] F. Lambiase, S.I. Scipioni, C.-J. Lee, D.-C. Ko, F. Liu, A State-of-the-Art Review on Advanced Joining Processes for Metal-Composite and Metal-Polymer Hybrid Structures, Materials (Basel) 14 (2021). https://doi.org/10.3390/ma14081890

[13] M. Merklein, M. Jäckisch, C.-M. Kuball, D. Römisch, S. Wiesenmayer, S. Wituschek, Mechanical joining of high-strength multi-material systems - trends and innovations, Mechanics & Industry 24 (2023). https://doi.org/10.1051/meca/2023013

[14] C. Wischer, W. Homberg, A contribution on versatile process chains: joining with adaptive joining elements, formed by friction spinning, Prod. Eng. Res. Devel. 16 (2022) 379-388. https://doi.org/10.1007/s11740-021-01094-8

[15] C. Wischer, W. Homberg, Further Development of an Adaptive Joining Technique Based on Friction Spinning to Produce Pre-Hole-Free Joints, KEM 926 (2022) 1468-1478. https://doi.org/10.4028/p-1n6741

[16] T. Borgert, M. Henke, W. Homberg, Investigations on the Influences of the Thermomechanical Manufacturing of Aluminium Auxiliary Joining Elements, JMMP 7 (2023) 147. https://doi.org/10.3390/jmmp7040147
# The continuous fibre injection process (CFIP): A novel approach to lightweight design of multi-material structural components

CRESCENTI Marc¹,ª*, RAFOLS Irene²,ᵇ, LARA Antoni²,ᶜ,
BEMANI Milad²,ᵈ, CASELLAS Daniel²,³,ᵉ

¹ Reinforce3D, Lligallo de Lorente 3, 43870 Amposta, Spain

² Eurecat, Centre Tecnològic de Catalunya, Product Innovation & Multiphysics Simulation Unit,
Universitat Autònoma, 23, 08290 Cerdanyola del Vallès, Spain

³ Luleå University of Technology, Division of Mechanics of Solid Materials, 971 87 Luleå,
Sweden

ª marc.crescenti@reinforce3d.com, ᵇ irene.rafols@eurecat.org, ᶜ toni.lara@eurecat.org,

ᵈ milad.bemani@eurecat.org, ᶜ daniel.casellas@eurecat.org

**Keywords:** Additive Manufacturing, Composites, Carbon Fiber, Joining Methods

**Abstract.** The combination of different materials enables to achieve highly efficient structures in terms of lightweight and mechanical performance, as well as in terms of manufacturing costs. However, the weakest points of these structures use to be the joints. For this reason, in the last years, many studies have dealt with joining technologies for dissimilar materials. The Reinforce3D's Continuous Fibre Injection Process (CFIP) technology delivers a unique method to join dissimilar materials. CFIP is based on injecting continuous fibers, such as carbon fibers, simultaneously with liquid resin into tubular cavities within the part. Then the resin is cured and the final composite part is obtained. This work focuses on the characterization of the mechanical properties of CFIP-made specimens and describes the potential lightweight benefits of the technology. Mechanical tests were performed under tensile and bending conditions following standardized methods. The lightweight potential is addressed by developing a representative case study by implementing finite element and topology optimization methods. The results of this case study were finally compared with a monomaterial equivalent component (aluminium) demonstrating the improvement that CFIP provides in terms of lightweight while keeping the strength.

**Introduction**

Materials and forming processes for lightweight construction of structural parts have been intensively investigated in the last decade. High strength steels and aluminum alloys, together with carbon fibre reinforced polymers are the typical material solutions, depending on the part design and expected costs. In the last years, the evolution of additive manufacturing (AM) technologies gives and additional strategy for weight saving thanks to their capability for manufacturing highly complex and optimized geometries, such as the ones obtained from topology optimization simulations. According to ref 1 and 2, topology optimized structures can reach around a 40% of weight reduction when compared to structures manufactured by conventional methods. Even higher lightweight ratios are possible when combining different materials, following the concept of *the right material at the right place*. Optimized multi-material construction requires efficient joining techniques. However, the weakest points of multi-material structures are usually the joints. Accordingly, many solutions based on adhesive or mechanical joining methods have been developed, with advantages but also some drawbacks that hampered their wide industrial application.

<figure><img src="image_1566.png" /></figure>
# The continuous fibre injection process (CFIP): A novel approach to lightweight design of multi-material structural components

CRESCENTI Marc¹,ª*, RAFOLS Irene²,ᵇ, LARA Antoni²,ᶜ,
BEMANI Milad²,ᵈ, CASELLAS Daniel²,³,ᵉ

¹ Reinforce3D, Lligallo de Lorente 3, 43870 Amposta, Spain

² Eurecat, Centre Tecnològic de Catalunya, Product Innovation & Multiphysics Simulation Unit,
Universitat Autònoma, 23, 08290 Cerdanyola del Vallès, Spain

³ Luleå University of Technology, Division of Mechanics of Solid Materials, 971 87 Luleå,
Sweden

ª marc.crescenti@reinforce3d.com, ᵇ irene.rafols@eurecat.org, ᶜ toni.lara@eurecat.org,

ᵈ milad.bemani@eurecat.org, ᶜ daniel.casellas@eurecat.org

**Keywords:** Additive Manufacturing, Composites, Carbon Fiber, Joining Methods

**Abstract.** The combination of different materials enables to achieve highly efficient structures in terms of lightweight and mechanical performance, as well as in terms of manufacturing costs. However, the weakest points of these structures use to be the joints. For this reason, in the last years, many studies have dealt with joining technologies for dissimilar materials. The Reinforce3D's Continuous Fibre Injection Process (CFIP) technology delivers a unique method to join dissimilar materials. CFIP is based on injecting continuous fibers, such as carbon fibers, simultaneously with liquid resin into tubular cavities within the part. Then the resin is cured and the final composite part is obtained. This work focuses on the characterization of the mechanical properties of CFIP-made specimens and describes the potential lightweight benefits of the technology. Mechanical tests were performed under tensile and bending conditions following standardized methods. The lightweight potential is addressed by developing a representative case study by implementing finite element and topology optimization methods. The results of this case study were finally compared with a monomaterial equivalent component (aluminium) demonstrating the improvement that CFIP provides in terms of lightweight while keeping the strength.

**Introduction**

Materials and forming processes for lightweight construction of structural parts have been intensively investigated in the last decade. High strength steels and aluminum alloys, together with carbon fibre reinforced polymers are the typical material solutions, depending on the part design and expected costs. In the last years, the evolution of additive manufacturing (AM) technologies gives and additional strategy for weight saving thanks to their capability for manufacturing highly complex and optimized geometries, such as the ones obtained from topology optimization simulations. According to ref 1 and 2, topology optimized structures can reach around a 40% of weight reduction when compared to structures manufactured by conventional methods. Even higher lightweight ratios are possible when combining different materials, following the concept of *the right material at the right place*. Optimized multi-material construction requires efficient joining techniques. However, the weakest points of multi-material structures are usually the joints. Accordingly, many solutions based on adhesive or mechanical joining methods have been developed, with advantages but also some drawbacks that hampered their wide industrial application.

<figure><img src="image_1567.png" /></figure>
An excellent joining approach would combine chemical and mechanical bonding and would give functionality to the structure. A novel approach based on applying continuous Carbon fibers (cCf) along the structure, joining dissimilar materials, and locally strengthening and/or increasing the stiffness has been recently developed [3,4]. This process is named Continuous Fiber Injection Process (CFIP). In CFIP, the cCfs are injected inside tubular cavities within a part. Three key principles enable the process: on one hand, the fibers are injected simultaneously with pressurized liquid resin, generating a flow that contributes (i) to drag the cCf along the tubular cavity and (ii) to lubricate it and, on the other hand, (iii) a pushing force is also applied on the fibers. After the injection, the resin is cured and solidified, bonding the fibers to the part so that the final composite part is obtained. Figure 1 describes this process when applied for manufacturing a car brake pedal.

Besides reinforcing, CFIP can also integrally join different parts by injecting the cCf along them, achieving fiber continuity from end-to-end and thus delivering an ultra-high joining performance, enabling the efficient manufacturing of large and multi-material/process structures. The brake pedal shown in Figure 1 is made by the integral joining of different parts selecting the most efficient material and manufacturing method at each zone according to mechanical requirements but also to cost a production targets. Particularly, the pedal shown in Figure 1 integrates the following materials and manufacturing technologies: an epoxy resin made by AM, a machined aluminum alloy and an extruded polyamide.

Consequently, CFIP delivers a unique method for the tailored reinforcement of multi-material parts with improved joint resistance.

<figure><img src="image_1567.png" /><figcaption>Figure 1: Scheme of the CFIP process: 1, part with tubular cavities; 2, cCf; 3, liquid resin; 4, injection of the composite material (cCfs + resin) inside the tubular cavities; and 5, final part.</figcaption></figure>

AM technologies allow the manufacturing of parts containing tubular cavities following complex trajectories in all directions. For this reason, when combined with AM, CFIP provides high freedom for placing the fibers within the part, enabling their alignment to the most efficient directions, even through printing layers, which use to be the weakest direction in AM parts. Parts made by any AM technology and material can be reinforced by CFIP, including rigid and flexible polymers manufactured by Selective Laser Sintering, Multi Jet Fusion or Stereolithography or metallic alloys processed by Selective Laser Melting.

While in the composite materials market CFIP is conceived as a new mouldless method for manufacturing highly complex and optimized geometries, in the AM market it is conceived as the first technology based on reinforcing the AM part after the printing process, instead of during it.
An excellent joining approach would combine chemical and mechanical bonding and would give functionality to the structure. A novel approach based on applying continuous Carbon fibers (cCf) along the structure, joining dissimilar materials, and locally strengthening and/or increasing the stiffness has been recently developed [3,4]. This process is named Continuous Fiber Injection Process (CFIP). In CFIP, the cCfs are injected inside tubular cavities within a part. Three key principles enable the process: on one hand, the fibers are injected simultaneously with pressurized liquid resin, generating a flow that contributes (i) to drag the cCf along the tubular cavity and (ii) to lubricate it and, on the other hand, (iii) a pushing force is also applied on the fibers. After the injection, the resin is cured and solidified, bonding the fibers to the part so that the final composite part is obtained. Figure 1 describes this process when applied for manufacturing a car brake pedal.

Besides reinforcing, CFIP can also integrally join different parts by injecting the cCf along them, achieving fiber continuity from end-to-end and thus delivering an ultra-high joining performance, enabling the efficient manufacturing of large and multi-material/process structures. The brake pedal shown in Figure 1 is made by the integral joining of different parts selecting the most efficient material and manufacturing method at each zone according to mechanical requirements but also to cost a production targets. Particularly, the pedal shown in Figure 1 integrates the following materials and manufacturing technologies: an epoxy resin made by AM, a machined aluminum alloy and an extruded polyamide.

Consequently, CFIP delivers a unique method for the tailored reinforcement of multi-material parts with improved joint resistance.

<figure><img src="image_1568.png" /><figcaption>Figure 1: Scheme of the CFIP process: 1, part with tubular cavities; 2, cCf; 3, liquid resin; 4, injection of the composite material (cCfs + resin) inside the tubular cavities; and 5, final part.</figcaption></figure>

AM technologies allow the manufacturing of parts containing tubular cavities following complex trajectories in all directions. For this reason, when combined with AM, CFIP provides high freedom for placing the fibers within the part, enabling their alignment to the most efficient directions, even through printing layers, which use to be the weakest direction in AM parts. Parts made by any AM technology and material can be reinforced by CFIP, including rigid and flexible polymers manufactured by Selective Laser Sintering, Multi Jet Fusion or Stereolithography or metallic alloys processed by Selective Laser Melting.

While in the composite materials market CFIP is conceived as a new mouldless method for manufacturing highly complex and optimized geometries, in the AM market it is conceived as the first technology based on reinforcing the AM part after the printing process, instead of during it.
This work shows the lightweight potential of the CFIP technology by comparing the weight and mechanical performance of two demonstrators: one topologically optimized and manufactured completely in aluminum by AM and the other developed by the reinforcement and integral joining by CFIP of two AM parts made of aluminium and polyamide.

## Material characterization

Specimens for tensile and bending tests were prepared using CFIP technology. The characteristics and dimensions of the CFIP specimens, for both the AM part and the injected composite material are given in Table 1. Figure 2 shows the test setup illustrating the direction of the injected composite material with respect to the loading direction. For the tensile specimens, a tab made of a fiberglass composite is added at the ends to give proper gripping during the tests and to promote the correct breakage of the samples (Figure 3).

<table><caption>Table 1: Characteristics of CFIP specimens.</caption><thead><tr><td>AM part</td><td>Injected composite inside the tubular cavity</td></tr></thead><tbody><tr><td>Tubular cavity dimensions: 2.5 x 12.4 mm</td><td>cCf and epoxy resin</td></tr><tr><td>Wall thickness: 0.5 mm</td><td>55% of cCf in volume</td></tr><tr><td>AM technology: Multi Jet Fusion</td><td></td></tr><tr><td>Material: Polyamide (PA) 12</td><td></td></tr></tbody></table>

<figure><img src="image_1568.png" /><figcaption>Figure 2: Tests setup. (a) Tensile specimens, with one cavity aligned to the loading direction. (b) Bending specimens, with two cavities perpendicularly oriented with respect to the applied load.</figcaption></figure>

Tensile tests were performed following the ASTM D3039 standard. Seven test repetitions were performed. All failed specimens show fiber breakage at the gauge zone. Results are shown in Table 2 and Figure 3a, which were calculated using the formulae described in the standard.

In a CFIP part, besides the AM material and the injected composite material, the interface between them also plays an important role. In order to evaluate the material properties of CFIP
This work shows the lightweight potential of the CFIP technology by comparing the weight and mechanical performance of two demonstrators: one topologically optimized and manufactured completely in aluminum by AM and the other developed by the reinforcement and integral joining by CFIP of two AM parts made of aluminium and polyamide.

## Material characterization

Specimens for tensile and bending tests were prepared using CFIP technology. The characteristics and dimensions of the CFIP specimens, for both the AM part and the injected composite material are given in Table 1. Figure 2 shows the test setup illustrating the direction of the injected composite material with respect to the loading direction. For the tensile specimens, a tab made of a fiberglass composite is added at the ends to give proper gripping during the tests and to promote the correct breakage of the samples (Figure 3).

<table><caption>Table 1: Characteristics of CFIP specimens.</caption><thead><tr><td>AM part</td><td>Injected composite inside the tubular cavity</td></tr></thead><tbody><tr><td>Tubular cavity dimensions: 2.5 x 12.4 mm</td><td>cCf and epoxy resin</td></tr><tr><td>Wall thickness: 0.5 mm</td><td>55% of cCf in volume</td></tr><tr><td>AM technology: Multi Jet Fusion</td><td></td></tr><tr><td>Material: Polyamide (PA) 12</td><td></td></tr></tbody></table>

<figure><img src="image_1569.png" /><figcaption>Figure 2: Tests setup. (a) Tensile specimens, with one cavity aligned to the loading direction. (b) Bending specimens, with two cavities perpendicularly oriented with respect to the applied load.</figcaption></figure>

Tensile tests were performed following the ASTM D3039 standard. Seven test repetitions were performed. All failed specimens show fiber breakage at the gauge zone. Results are shown in Table 2 and Figure 3a, which were calculated using the formulae described in the standard.

In a CFIP part, besides the AM material and the injected composite material, the interface between them also plays an important role. In order to evaluate the material properties of CFIP
samples in a setup where the interface transfers significant loads, sandwich-like specimens under bending load have been tested following the ASTM D790 standard. Five test repetitions were performed. The load vs displacement curves are shown in Figure 3b, and the results are summarized in Table 2, which were calculated using the formulae described in the standard. Specimens 1A and 1C experienced the breakage of the upper fibers at compression followed immediately by the failure of the upper interface. Specimens 1B, 1D and 1E experienced the breakage of the upper fibers at compression followed by the breakage of the lower fibers at tension after a relevant deformation. Thus, in all the specimens the interface could withstand the maximum load and the fibres could be brought to their limit, demonstrating that CFIP can also be an efficient solution in cases where the interface withstands relevant stresses.

<figure><img src="image_1569.png" /><figcaption>Figure 3: Test setups: (a) image of a broken specimen, and stress vs strain curve showing results for 7 tensile specimens; (b) image of broken specimens, and stress vs strain curve showing results for 5 bending specimens.</figcaption></figure>

For simulation and analysis purposes, it is needed to know the mechanical properties of the injected composite material inside the cavity. These results have been calculated from the tensile test results subtracting the area of the wall from the total transverse area according to the measured wall thickness. Results are included in Table 2.

A part of the CFIP technology demonstrator and the whole AM demonstrator shown in the next section, are made of AlSi10Mg alloy manufactured by Laser powder bed fusion (PBF-LB/M). Thus, 3 dog bone specimens were manufactured with the same processing parameters and tested following the EN ISO 6892-1 standard. The obtained tensile properties are included in Table 2, which are similar to previously reported values for similar Al alloys processed by PBF-LB/M [5].
samples in a setup where the interface transfers significant loads, sandwich-like specimens under bending load have been tested following the ASTM D790 standard. Five test repetitions were performed. The load vs displacement curves are shown in Figure 3b, and the results are summarized in Table 2, which were calculated using the formulae described in the standard. Specimens 1A and 1C experienced the breakage of the upper fibers at compression followed immediately by the failure of the upper interface. Specimens 1B, 1D and 1E experienced the breakage of the upper fibers at compression followed by the breakage of the lower fibers at tension after a relevant deformation. Thus, in all the specimens the interface could withstand the maximum load and the fibres could be brought to their limit, demonstrating that CFIP can also be an efficient solution in cases where the interface withstands relevant stresses.

<figure><img src="image_1570.png" /><figcaption>Figure 3: Test setups: (a) image of a broken specimen, and stress vs strain curve showing results for 7 tensile specimens; (b) image of broken specimens, and stress vs strain curve showing results for 5 bending specimens.</figcaption></figure>

For simulation and analysis purposes, it is needed to know the mechanical properties of the injected composite material inside the cavity. These results have been calculated from the tensile test results subtracting the area of the wall from the total transverse area according to the measured wall thickness. Results are included in Table 2.

A part of the CFIP technology demonstrator and the whole AM demonstrator shown in the next section, are made of AlSi10Mg alloy manufactured by Laser powder bed fusion (PBF-LB/M). Thus, 3 dog bone specimens were manufactured with the same processing parameters and tested following the EN ISO 6892-1 standard. The obtained tensile properties are included in Table 2, which are similar to previously reported values for similar Al alloys processed by PBF-LB/M [5].
<table><caption>Table 2: Mechanical properties obtained for the CFIP material, the injected composite inside the cavities of CFIP specimens, and the Al alloy used for the AM demonstrator. CV is the coefficient of variance, calculated as the ratio between the standard deviation and the mean value.</caption><thead><tr><th>Material</th><th>Test</th><th>Property</th><th>Mean value</th><th>CV (%)</th></tr></thead><tbody><tr><td rowspan="6">CFIP (AM part + injected composite)</td><td rowspan="3">Tensile</td><td>Young's modulus (GPa)</td><td>80.2</td><td>3.8</td></tr><tr><td>Tensile strength (MPa)</td><td>1120</td><td>6.3</td></tr><tr><td>Elongation at fracture (%)</td><td>1.45</td><td>4.7</td></tr><tr><td rowspan="2">Bending</td><td>Bending strength (MPa)</td><td>422</td><td>5.8</td></tr><tr><td>Flexural stiffness (GPa)</td><td>42.9</td><td>2.6</td></tr><tr><td rowspan="2">Injected composite (cCf + epoxy resin)</td><td rowspan="2">Tensile</td><td>Young's modulus (GPa)</td><td>120.6</td><td>-</td></tr><tr><td>Tensile strength (MPa)</td><td>1683</td><td>-</td></tr><tr><td rowspan="3">AlSi10Mg manufactured by AM</td><td rowspan="3">Tensile</td><td>Young's modulus (GPa)</td><td>68.4</td><td>0.1</td></tr><tr><td>Yield strength (MPa)</td><td>325</td><td>0.3</td></tr><tr><td>Tensile strength (MPa)</td><td>465</td><td>0.2</td></tr></tbody></table>

The common way to compare the lightweight potential of different materials is through the ratio of strength versus density, known as the specific strength. Such value was calculated for the studied CFIP material (density of 1.22 g/cm³), and compared with the AlSi10Mg manufactured by AM (density of 2.7 g/cm³), the polyamide 12 (PA12) manufactured by AM (strength of 48 MPa and density of 1.01 g/cm³, extracted from ref 6), and other lightweight solutions based on cCfs manufactured by AM available in the literature (tensile strength of 700 MPa and density of 1.19 g/cm³, extracted from ref 7 and 8). The results are shown in Figure 4, where it is observed that the specific strength of CFIP material is clearly higher than the other materials and processing methods. Therefore, CFIP stands as the technology with the highest lightweight capacity while also allowing to take advantage of the freedom for design given by the flexibility of the AM techniques.

<figure><img src="image_1570.png" /><figcaption>Figure 4: Specific strength of the different material options to build lightweight parts by AM: CFIP material, cCf composite material [7,8], Al alloy and polyamide 12 [6].</figcaption></figure>
<table><caption>Table 2: Mechanical properties obtained for the CFIP material, the injected composite inside the cavities of CFIP specimens, and the Al alloy used for the AM demonstrator. CV is the coefficient of variance, calculated as the ratio between the standard deviation and the mean value.</caption><thead><tr><th>Material</th><th>Test</th><th>Property</th><th>Mean value</th><th>CV (%)</th></tr></thead><tbody><tr><td rowspan="6">CFIP (AM part + injected composite)</td><td rowspan="3">Tensile</td><td>Young's modulus (GPa)</td><td>80.2</td><td>3.8</td></tr><tr><td>Tensile strength (MPa)</td><td>1120</td><td>6.3</td></tr><tr><td>Elongation at fracture (%)</td><td>1.45</td><td>4.7</td></tr><tr><td rowspan="2">Bending</td><td>Bending strength (MPa)</td><td>422</td><td>5.8</td></tr><tr><td>Flexural stiffness (GPa)</td><td>42.9</td><td>2.6</td></tr><tr><td rowspan="2">Injected composite (cCf + epoxy resin)</td><td rowspan="2">Tensile</td><td>Young's modulus (GPa)</td><td>120.6</td><td>-</td></tr><tr><td>Tensile strength (MPa)</td><td>1683</td><td>-</td></tr><tr><td rowspan="3">AlSi10Mg manufactured by AM</td><td rowspan="3">Tensile</td><td>Young's modulus (GPa)</td><td>68.4</td><td>0.1</td></tr><tr><td>Yield strength (MPa)</td><td>325</td><td>0.3</td></tr><tr><td>Tensile strength (MPa)</td><td>465</td><td>0.2</td></tr></tbody></table>

The common way to compare the lightweight potential of different materials is through the ratio of strength versus density, known as the specific strength. Such value was calculated for the studied CFIP material (density of 1.22 g/cm³), and compared with the AlSi10Mg manufactured by AM (density of 2.7 g/cm³), the polyamide 12 (PA12) manufactured by AM (strength of 48 MPa and density of 1.01 g/cm³, extracted from ref 6), and other lightweight solutions based on cCfs manufactured by AM available in the literature (tensile strength of 700 MPa and density of 1.19 g/cm³, extracted from ref 7 and 8). The results are shown in Figure 4, where it is observed that the specific strength of CFIP material is clearly higher than the other materials and processing methods. Therefore, CFIP stands as the technology with the highest lightweight capacity while also allowing to take advantage of the freedom for design given by the flexibility of the AM techniques.

<figure><img src="image_1571.png" /><figcaption>Figure 4: Specific strength of the different material options to build lightweight parts by AM: CFIP material, cCf composite material [7,8], Al alloy and polyamide 12 [6].</figcaption></figure>
## CFIP technology validation: case study

The mechanical and lightweight performance of a structure is derived from the materials properties and geometry. To evaluate the performance of CFIP at the demonstrator level, a representative case study has been defined, consisting of a structure of 120x100x240 mm clamped through 4 bolts and submitted to a load in the Y direction of -5000 N. Figure 5 shows the created design space highlighting the clamping and loading zones. The objective is to minimize the weight while keeping the structural integrity following two different strategies:

(a) Demonstrator designed to fully exploit the lightweight potential of the CFIP technology. From now on it is named as CFIP demonstrator.

(b) Demonstrator manufactured with AlSi10Mg alloy by Laser powder bed fusion (PBF-LB/M) with a topologically optimized geometry to fully exploit its lightweight potential. It is named as AM demonstrator.

<figure><img src="image_1571.png" /><figcaption>Figure 5: Design space of the defined case study. Clamped zones in red and loading zone in blue.</figcaption></figure>

The CFIP demonstrator has been developed by the reinforcement and integral joining of two different parts manufactured by AM (Figure 6). At the clamped zone and specially around the bolts and CFIP units, high stress levels in all directions are produced, so an isotropic material with high specific properties is the best option. Thus, an AlSi10Mg alloy part manufactured by PBF-LB/M has been used at this zone, aiming to efficiently transfer the loads from the bolts to the injected CFIP units. For the central and loading zones, a PA12 part made by Multi Jet Fusion has been used aiming to save costs while providing the required mechanical properties.

The design of the fibre trajectories has been done by following three main criteria: (i) trace them from the loading zone to the clamped zone in order to efficiently transfer the loads, (ii) separate them where the moments are higher in order to increase the inertia and strength, and (iii) connect both parts in order to integrally join them.

Six CFIP units made of cCfs and epoxy resin have been used. A variant cross-section from circular (clamped zone) to squared (loading zone) has been designed with an effective diameter of 7 mm, aiming to efficiently transfer the shear loads between the injected composite material and the AM parts at both regions.
## CFIP technology validation: case study

The mechanical and lightweight performance of a structure is derived from the materials properties and geometry. To evaluate the performance of CFIP at the demonstrator level, a representative case study has been defined, consisting of a structure of 120x100x240 mm clamped through 4 bolts and submitted to a load in the Y direction of -5000 N. Figure 5 shows the created design space highlighting the clamping and loading zones. The objective is to minimize the weight while keeping the structural integrity following two different strategies:

(a) Demonstrator designed to fully exploit the lightweight potential of the CFIP technology. From now on it is named as CFIP demonstrator.

(b) Demonstrator manufactured with AlSi10Mg alloy by Laser powder bed fusion (PBF-LB/M) with a topologically optimized geometry to fully exploit its lightweight potential. It is named as AM demonstrator.

<figure><img src="image_1572.png" /><figcaption>Figure 5: Design space of the defined case study. Clamped zones in red and loading zone in blue.</figcaption></figure>

The CFIP demonstrator has been developed by the reinforcement and integral joining of two different parts manufactured by AM (Figure 6). At the clamped zone and specially around the bolts and CFIP units, high stress levels in all directions are produced, so an isotropic material with high specific properties is the best option. Thus, an AlSi10Mg alloy part manufactured by PBF-LB/M has been used at this zone, aiming to efficiently transfer the loads from the bolts to the injected CFIP units. For the central and loading zones, a PA12 part made by Multi Jet Fusion has been used aiming to save costs while providing the required mechanical properties.

The design of the fibre trajectories has been done by following three main criteria: (i) trace them from the loading zone to the clamped zone in order to efficiently transfer the loads, (ii) separate them where the moments are higher in order to increase the inertia and strength, and (iii) connect both parts in order to integrally join them.

Six CFIP units made of cCfs and epoxy resin have been used. A variant cross-section from circular (clamped zone) to squared (loading zone) has been designed with an effective diameter of 7 mm, aiming to efficiently transfer the shear loads between the injected composite material and the AM parts at both regions.
<figure><img src="image_1572.png" /><figcaption>Figure 6. Design of the CFIP demonstrator: AlSi10Mg alloy part in grey, PA12 part in cyan, and injected composite material by CFIP in black.</figcaption></figure>

The mechanical performance of this design was analyzed by the Finite Element Method (FEM) in ANSYS. The AlSi10Mg and PA12 materials have been modelled by defining a plasticity material model based on a bilinear isotropic law. The CFIP units have been modeled as a unidirectional composite material using an orthotropic model. Discrete orientations have been also defined on the CFIP units to align the X direction of the material to the longitudinal direction of the units at each element. The geometry has been meshed with tetrahedral quadratic elements with an average size of 2 mm, as shown in Figure 7a. The obtained displacements after FEM analysis are shown in Figure 7b.

<figure><img src="image_1573.png" /><figcaption>Figure 7. FEM model of the CFIP demonstrator: (a) Mesh, (b) Displacements (mm).</figcaption></figure>

The stress distribution in the CFIP demonstrator must be analyzed in detail in the AM parts, the injected composite material and also at the interface itself. Figure 8 shows the FEM results for such regions. In the aluminum part, the stresses are slightly above the yield strength at the clamped zones but below the tensile strength (Figure 8a), indicating that very localized plasticity could occur. In the PA12 part the stresses are far below the yield strength (Figure 8b). Figure 8c shows the distribution of shear stresses at the interface. It is observed that the stresses reach significant levels above 20 MPa only in localized zones close to the clamped regions. The stresses in the injected composite material have also been analyzed, showing values much below the tensile strength. A buckling analysis has been also performed on the structure, indicating that no global or local buckling will occur.
<figure><img src="image_1573.png" /><figcaption>Figure 6. Design of the CFIP demonstrator: AlSi10Mg alloy part in grey, PA12 part in cyan, and injected composite material by CFIP in black.</figcaption></figure>

The mechanical performance of this design was analyzed by the Finite Element Method (FEM) in ANSYS. The AlSi10Mg and PA12 materials have been modelled by defining a plasticity material model based on a bilinear isotropic law. The CFIP units have been modeled as a unidirectional composite material using an orthotropic model. Discrete orientations have been also defined on the CFIP units to align the X direction of the material to the longitudinal direction of the units at each element. The geometry has been meshed with tetrahedral quadratic elements with an average size of 2 mm, as shown in Figure 7a. The obtained displacements after FEM analysis are shown in Figure 7b.

<figure><img src="image_1574.png" /><figcaption>Figure 7. FEM model of the CFIP demonstrator: (a) Mesh, (b) Displacements (mm).</figcaption></figure>

The stress distribution in the CFIP demonstrator must be analyzed in detail in the AM parts, the injected composite material and also at the interface itself. Figure 8 shows the FEM results for such regions. In the aluminum part, the stresses are slightly above the yield strength at the clamped zones but below the tensile strength (Figure 8a), indicating that very localized plasticity could occur. In the PA12 part the stresses are far below the yield strength (Figure 8b). Figure 8c shows the distribution of shear stresses at the interface. It is observed that the stresses reach significant levels above 20 MPa only in localized zones close to the clamped regions. The stresses in the injected composite material have also been analyzed, showing values much below the tensile strength. A buckling analysis has been also performed on the structure, indicating that no global or local buckling will occur.
<figure><img src="image_1574.png" /><figcaption>Figure 8: FEM results on the CFIP demonstrator: (a) Von-Mises stresses (MPa) at the AlSi10Mg alloy part. (b) Von-Mises stresses (MPa) at the PA12 part. (c) Longitudinal stresses (MPa) at the injected composite material. (d) Shear stresses (MPa) at the interface.</figcaption></figure>

The AM demonstrator has been topology optimized by using the ALTAIR Optistruct solver, aiming to fully exploit its lightweight potential. The yield strength of the AlSi10Mg material has been defined as a constraint and the objective has been set to minimize the weight. The obtained geometry has been detailed in a CAD software in order to reach a functional design. FEM analysis have been also performed on this design using ANSYS. A plasticity material model has been defined by using a bilinear isotropic law. The geometry was meshed using quadratic tetrahedral elements with an average size of 2 mm (Figure 9a). A buckling analysis was also performed on this demonstrator, indicating that no global or local buckling will occur. Figure 9 also shows the distribution of displacements and stresses. It is observed how the stresses at the clamped zones are just at the yield strength of the material (Figure 9c), a really similar situation as in the CFIP demonstrator. It is also observed that the displacements are significantly lower compared to the CFIP structure, although they are relatively low in both cases compared to the dimensions of the structure and no requirements in terms of stiffness have been defined in the case study.

The final weight of each demonstrator is shown in Table 3. CFIP technology gives a 32% of weight reduction compared to the topologically optimized AM design in aluminum. As shown in other works [1, 2], such topology optimization for AM allows a 40% of weight savings with respect to conventionally manufactured parts. It means that the lightweight potential of CFIP technology can reach up to 60% when compared to traditional manufacturing methods, thanks to the efficient combination of different materials in highly optimized designs.
<figure><img src="image_1575.png" /><figcaption>Figure 8: FEM results on the CFIP demonstrator: (a) Von-Mises stresses (MPa) at the AlSi10Mg alloy part. (b) Von-Mises stresses (MPa) at the PA12 part. (c) Longitudinal stresses (MPa) at the injected composite material. (d) Shear stresses (MPa) at the interface.</figcaption></figure>

The AM demonstrator has been topology optimized by using the ALTAIR Optistruct solver, aiming to fully exploit its lightweight potential. The yield strength of the AlSi10Mg material has been defined as a constraint and the objective has been set to minimize the weight. The obtained geometry has been detailed in a CAD software in order to reach a functional design. FEM analysis have been also performed on this design using ANSYS. A plasticity material model has been defined by using a bilinear isotropic law. The geometry was meshed using quadratic tetrahedral elements with an average size of 2 mm (Figure 9a). A buckling analysis was also performed on this demonstrator, indicating that no global or local buckling will occur. Figure 9 also shows the distribution of displacements and stresses. It is observed how the stresses at the clamped zones are just at the yield strength of the material (Figure 9c), a really similar situation as in the CFIP demonstrator. It is also observed that the displacements are significantly lower compared to the CFIP structure, although they are relatively low in both cases compared to the dimensions of the structure and no requirements in terms of stiffness have been defined in the case study.

The final weight of each demonstrator is shown in Table 3. CFIP technology gives a 32% of weight reduction compared to the topologically optimized AM design in aluminum. As shown in other works [1, 2], such topology optimization for AM allows a 40% of weight savings with respect to conventionally manufactured parts. It means that the lightweight potential of CFIP technology can reach up to 60% when compared to traditional manufacturing methods, thanks to the efficient combination of different materials in highly optimized designs.
<figure><img src="image_1575.png" /><figcaption>Figure 9: (a) Topology optimized geometry for AM demonstrator showing the mesh for FEM analysis. FEM results: (b) Displacement (mm), (c) Von-Mises stresses (MPa).</figcaption></figure>

<table><caption>Table 3: Comparison of the weight of the two demonstrators.</caption><thead><tr><td>Demonstrator</td><td>Weight (g)</td><td>Weight reduction (%)</td></tr></thead><tbody><tr><td>Topologically optimized design for AM, Al alloy</td><td>329</td><td>-</td></tr><tr><td>CFIP technology, multi-material solution</td><td>224</td><td>32</td></tr></tbody></table>

## Conclusions

This work presents the CFIP technology as an innovative approach to extremely optimize the lightweight and mechanical performance of structures, thanks to the reinforcement of highly optimized parts with continuous carbon fibers. The performance was evaluated at material and at demonstrator level for the CFIP approach and for AM in aluminum alloy.

The development of the CFIP demonstrator was assessed by the reinforcement and integral joining of two parts made by AM, one in aluminum alloy and the other in polyamide 12. The design was numerically evaluated by the FEM and compared with a topologically optimized part developed completely by AM in aluminum alloy. A weight reduction of 32% was obtained.

Such significant lightweight potential is attained thanks to the use of highly optimized AM parts, the efficient combination of materials and the optimal distribution of the continuous carbon fibers within the AM parts, which could not be achieved by other composites manufacturing methods based on layer-by-layer distribution of the fibers. Compared to conventionally manufactured components, the CFIP technology can reach a weight reduction of 60%.
<figure><img src="image_1576.png" /><figcaption>Figure 9: (a) Topology optimized geometry for AM demonstrator showing the mesh for FEM analysis. FEM results: (b) Displacement (mm), (c) Von-Mises stresses (MPa).</figcaption></figure>

<table><caption>Table 3: Comparison of the weight of the two demonstrators.</caption><thead><tr><td>Demonstrator</td><td>Weight (g)</td><td>Weight reduction (%)</td></tr></thead><tbody><tr><td>Topologically optimized design for AM, Al alloy</td><td>329</td><td>-</td></tr><tr><td>CFIP technology, multi-material solution</td><td>224</td><td>32</td></tr></tbody></table>

## Conclusions

This work presents the CFIP technology as an innovative approach to extremely optimize the lightweight and mechanical performance of structures, thanks to the reinforcement of highly optimized parts with continuous carbon fibers. The performance was evaluated at material and at demonstrator level for the CFIP approach and for AM in aluminum alloy.

The development of the CFIP demonstrator was assessed by the reinforcement and integral joining of two parts made by AM, one in aluminum alloy and the other in polyamide 12. The design was numerically evaluated by the FEM and compared with a topologically optimized part developed completely by AM in aluminum alloy. A weight reduction of 32% was obtained.

Such significant lightweight potential is attained thanks to the use of highly optimized AM parts, the efficient combination of materials and the optimal distribution of the continuous carbon fibers within the AM parts, which could not be achieved by other composites manufacturing methods based on layer-by-layer distribution of the fibers. Compared to conventionally manufactured components, the CFIP technology can reach a weight reduction of 60%.
## Acknowledgements

This work was partially funded by the Catalan Government through the funding grant ACCIÓ-Eurecat (Project TRAÇA-M3DTALL). M.B. acknowledges the funding from the European Union's Horizon 2020 research and innovation program under the Marie Skłodowska-Curie grant agreement No 101034328. This paper reflects only the author's view and the Research Executive Agency, and the European Commission are not responsible for any use that may be made of the information it contains.

## References

[1] A. Merulla, A. Gatto, E. Bassoli, S. I. Munteanu, B. Gheorghiu, M. A. Pop, T. Bedo, D. Munteanu, Weight reduction by topology optimization of an engine subframe mount, designed for additive manufacturing production, Materials Today: Proceedings 19 (2019): 1014-1018. https://doi.org/10.1016/j.matpr.2019.08.015

[2] EOS. Innovation story: RUAG 3D printed satellite parts. https://www.eos.info/en/innovations/all-3d-printing-applications/aerospace/aerospace-case-studies/ruag-aerospace-3d-printed-satellite-components

[3] M Crescenti, J.M. Lluís, Patent WO2016092132A1. (2014)

[4] Information on https://www.youtube.com/watch?v=c-ffXIMG8FE (2023)

[5] J. Gong, K. Wei, M. Liu, W. Song, X. Li, X. Zeng, Microstructure and mechanical properties of AlSi10Mg alloy built by laser powder bed fusion/direct energy deposition hybrid laser additive manufacturing, Additive Manufacturing 59 (2022) 103160. https://doi.org/10.1016/j.addma.2022.103160

[6] HP 3D High Reusability PA 12. Material datasheet (2018).

[7] J. Justo, L. Távara, L. García-Guzmán, F. París. Characterization of 3D printed long fibre reinforced composites. Composite Structures 185 (2018) 537-548. https://doi.org/10.1016/j.compstruct.2017.11.052

[8] A. Todoroki, T. Oasada, S. Mizutani, Y. Suzuki, M. Ueda, R. Matsuzaki, Y. Hirano, Tensile property evaluations of 3D printed continuous carbon fiber reinforced thermoplastic composites, Advanced Composite Materials, 29 (2020) 147-162. https://doi.org/10.1080/09243046.2019.1650323
## Acknowledgements

This work was partially funded by the Catalan Government through the funding grant ACCIÓ-Eurecat (Project TRAÇA-M3DTALL). M.B. acknowledges the funding from the European Union's Horizon 2020 research and innovation program under the Marie Skłodowska-Curie grant agreement No 101034328. This paper reflects only the author's view and the Research Executive Agency, and the European Commission are not responsible for any use that may be made of the information it contains.

## References

[1] A. Merulla, A. Gatto, E. Bassoli, S. I. Munteanu, B. Gheorghiu, M. A. Pop, T. Bedo, D. Munteanu, Weight reduction by topology optimization of an engine subframe mount, designed for additive manufacturing production, Materials Today: Proceedings 19 (2019): 1014-1018. https://doi.org/10.1016/j.matpr.2019.08.015

[2] EOS. Innovation story: RUAG 3D printed satellite parts. https://www.eos.info/en/innovations/all-3d-printing-applications/aerospace/aerospace-case-studies/ruag-aerospace-3d-printed-satellite-components

[3] M Crescenti, J.M. Lluís, Patent WO2016092132A1. (2014)

[4] Information on https://www.youtube.com/watch?v=c-ffXIMG8FE (2023)

[5] J. Gong, K. Wei, M. Liu, W. Song, X. Li, X. Zeng, Microstructure and mechanical properties of AlSi10Mg alloy built by laser powder bed fusion/direct energy deposition hybrid laser additive manufacturing, Additive Manufacturing 59 (2022) 103160. https://doi.org/10.1016/j.addma.2022.103160

[6] HP 3D High Reusability PA 12. Material datasheet (2018).

[7] J. Justo, L. Távara, L. García-Guzmán, F. París. Characterization of 3D printed long fibre reinforced composites. Composite Structures 185 (2018) 537-548. https://doi.org/10.1016/j.compstruct.2017.11.052

[8] A. Todoroki, T. Oasada, S. Mizutani, Y. Suzuki, M. Ueda, R. Matsuzaki, Y. Hirano, Tensile property evaluations of 3D printed continuous carbon fiber reinforced thermoplastic composites, Advanced Composite Materials, 29 (2020) 147-162. https://doi.org/10.1080/09243046.2019.1650323
# Forming of mechanically interlocked aluminium and carbon fibre reinforced polymer parts with complex geometry

LATORRE Núria¹,²,a *, CASELLAS Daniel¹,³,b, COSTA Josep²,c,
GARCIA-LLAMAS Eduard¹,d and PUJANTE Jaume¹,e

¹Eurecat, Technological Centre of Catalonia, Parc Tecnològic del Vallès, Av. Universitat Autònoma, 23 - 08290 Cerdanyola del Vallès, Spain

²AMADE Research Group, Polytechnic School, University of Girona, Campus Montilivi s/n, E-17003 Girona, Spain

³List Division of Mechanics of Solid Materials, Luleå University of Technology, 971 87 Luleå, Sweden

a nuria.latorre@eurecat.org, b daniel.casellas@eurecat.org, c josep.costa@udg.edu, d eduard.garcia@eurecat.org, e jaume.pujante@eurecat.org

**Keywords:** Metal, CFRP, Punching, Stamping, Automotive

**Abstract.** Forming of aluminium-CFRP hybrid structures into complex shapes is key to decrease environmental impact in automotive industry. However, challenges such as preserving joint integrity after forming operations must be assessed. Therefore, the authors of this work have cold stamped hybrid aluminium-CFRP panels into omega shaped profiles with and without a mechanical interlocking joining technology. The effect of lubricant application, of the CFRP positioning (inside or outside the omega profile), and of the number of mechanical joints were studied. It was concluded that it is possible to cold stamp aluminium-CFRP prepreg panels even with mechanical joints into complex profiles when lubricant is used. Moreover, the position of the CFRP prepreg has a strong impact on the flange springback of the stamped part.

**Introduction**

Nowadays, environmental demands are pushing automotive manufacturers to reduce vehicle weight without compromising its performance and safety. An interesting strategy to achieve that is multi-material design of structural parts of vehicle's components, giving a crucial role to the joining between dissimilar lightweight materials [1].

However, conventional joining strategies between metals and Carbon Fibre Reinforced Polymers (CFRP) either require drilling a hole in the composite material, which can lead to damages causing a decrease on the load bearing capacity of the components, or the increase of the weight of the part with the incorporation of fasteners [2-6]. Alternative mechanical joining strategies have been developed to overcome these drawbacks, but they involve complex and costly processing [7].

This led the authors of this work to develop a mechanical joining strategy between aluminium and CFRP consisting of a single-step punching process. Such process is simple, does not perforate the composite material, nor incorporate any extra weight on the part, and can be easily implemented in the automotive manufacturing lines [8]. Such mechanical interlocking joint improved the shear strength of the co-cured joint between 41 % and 57 %, and the absorbed energy between 94 % and 205 % for the studied materials (Fig. 1).

<figure><img src="image_1576.png" /></figure>
# Forming of mechanically interlocked aluminium and carbon fibre reinforced polymer parts with complex geometry

LATORRE Núria¹,²,a *, CASELLAS Daniel¹,³,b, COSTA Josep²,c,
GARCIA-LLAMAS Eduard¹,d and PUJANTE Jaume¹,e

¹Eurecat, Technological Centre of Catalonia, Parc Tecnològic del Vallès, Av. Universitat Autònoma, 23 - 08290 Cerdanyola del Vallès, Spain

²AMADE Research Group, Polytechnic School, University of Girona, Campus Montilivi s/n, E-17003 Girona, Spain

³List Division of Mechanics of Solid Materials, Luleå University of Technology, 971 87 Luleå, Sweden

a nuria.latorre@eurecat.org, b daniel.casellas@eurecat.org, c josep.costa@udg.edu,

d eduard.garcia@eurecat.org, e jaume.pujante@eurecat.org

**Keywords:** Metal, CFRP, Punching, Stamping, Automotive

**Abstract.** Forming of aluminium-CFRP hybrid structures into complex shapes is key to decrease environmental impact in automotive industry. However, challenges such as preserving joint integrity after forming operations must be assessed. Therefore, the authors of this work have cold stamped hybrid aluminium-CFRP panels into omega shaped profiles with and without a mechanical interlocking joining technology. The effect of lubricant application, of the CFRP positioning (inside or outside the omega profile), and of the number of mechanical joints were studied. It was concluded that it is possible to cold stamp aluminium-CFRP prepreg panels even with mechanical joints into complex profiles when lubricant is used. Moreover, the position of the CFRP prepreg has a strong impact on the flange springback of the stamped part.

**Introduction**

Nowadays, environmental demands are pushing automotive manufacturers to reduce vehicle weight without compromising its performance and safety. An interesting strategy to achieve that is multi-material design of structural parts of vehicle's components, giving a crucial role to the joining between dissimilar lightweight materials [1].

However, conventional joining strategies between metals and Carbon Fibre Reinforced Polymers (CFRP) either require drilling a hole in the composite material, which can lead to damages causing a decrease on the load bearing capacity of the components, or the increase of the weight of the part with the incorporation of fasteners [2-6]. Alternative mechanical joining strategies have been developed to overcome these drawbacks, but they involve complex and costly processing [7].

This led the authors of this work to develop a mechanical joining strategy between aluminium and CFRP consisting of a single-step punching process. Such process is simple, does not perforate the composite material, nor incorporate any extra weight on the part, and can be easily implemented in the automotive manufacturing lines [8]. Such mechanical interlocking joint improved the shear strength of the co-cured joint between 41 % and 57 %, and the absorbed energy between 94 % and 205 % for the studied materials (Fig. 1).

<figure><img src="image_1577.png" /></figure>