## 2 Materials and Experimental Work
### Finite Element Modelling
For the damage evolution criterion, \(G_{n},\ G_{s},\ G_{s}\), and \(G_{t}\) represent the work done by the tractions and their conjugate relative displacements in the normal, first, and second shear directions, respectively, and \(G_{T}\) is the total energy release rate. Moreover, \(\delta_{m}^{f}\) and \(\delta_{m}^{0}\) specify the effective displacement at complete failure and that relative to the effective displacement at the initiation of damage, respectively; \(G^{c}\) is the energy dissipated due to failure; and \(\eta\) is a material parameter [35]. The adhesion strength between the cover and core layers decreased sharply with the increasing temperature, as the higher temperatures increased the overall ductility of the adhesive, making crack propagation much smoother [36]. The properties of the cohesive  
Figure 6: Empirical stress–strain curves for TS275 with the Swift law. Figure 7: Illustration of mixed-mode response of cohesive elements [35]. elements in this study were defined based on the results presented in previous studies [37; 38], in which the peak traction stresses in the normal and shear modes were reduced by 83% each when the temperature increased from RT to 110 \({}^{\circ}\)C for an epoxy-based adhesive. The aim was to replicate and validate the model, where after the deep-drawing experiments, the adhesion was partially deteriorated. In a perfect case, the parameters for such cohesive zone modelling must be determined for the current material combinations; this point is currently under investigation in the framework of the follow-up DFG project, number 330043166. In this study, the deep-drawing behaviour of FML was simulated below the melting point of the PA6 matrix in the organosheet RG, i.e., 200 \({}^{\circ}\)C. The material properties for the cohesive elements in the model are summarized in Table 6. The size of the cohesive element was set to 1 mm and the thickness to 15 \(\upmu\)m. As the adhesion promotor was based on PA6, the density of the cohesive layer was set to be equal to PA6, i.e., 1.08 g/mm\({}^{3}\). In addition, the interaction between the FML cover and core layer could also be assumed to follow a tangential friction behaviour, taking into account the possible relative sliding between the cover and core layers at 200 \({}^{\circ}\)C, which was similar to the relative sliding that occurred in thermoplastic-based FML composed of Al/GF-PA6/Al when the melting point of the core layer PA6 was exceeded, i.e., the friction coefficient between Al and PA6 varied between 0.1 and 0.9 at temperatures above 220 \({}^{\circ}\)C [34]. This is influenced by temperature, pressure, and sliding velocity. In this study, Columbian friction coefficients \(\mu\) of 0.5 and 0.9 were considered to define the frictional behaviour in the model. Finally, a tie constraint was applied between the cover and core layers, assuming no relative sliding between the cover and core layers at 200 \({}^{\circ}\)C. The simulation results of different interaction conditions of the layers are compared with the experimental results at the end of this paper. The different conditions for defining the contact behaviour between the cover and core layers are summarized in Table 7. Similarly, the damage fracture behaviour of the RG core was not considered in the current case. The dimension of the round sample in the simulation model was 180 mm in diameter. The upper surface of the die in the model was fixed, and a holding force was applied to the lower surface of the holder. The sample was compressed by the holder under this load, and once the punch began to move, the sample was able to slide inwards between the die and the holder (see Figure 8).