## 5 Numerical and Experimental Results
\begin{table}
\begin{tabular}{l c c} \hline \hline
**Source** & \multicolumn{2}{c}{**Not Reinforced Sample [kN]**} \\ \hline Experimental Sample & & 39.5 \\ LS-DTNA Model & First Peak [kN] & Second Peak [kN] & ERR [\%] (Max Peak) \\ MAT58\_MAT63 & 52.7 & - & 25.1 \\ \hline \hline \end{tabular}
\end{table} 
Table 9: Not reinforced maximum load comparison. internal contact load exactly follows the global failure load of the simple specimen. The absence of steel reinforcement increases the localization of the damage and the initiation of the bearing fracture mode in the contact region, as in Figure 11, but no evidence of a propagation of such mode of fracture is determined, as revealed in the experimental test shown in Figure 5B. This is probably due to the compressive guided mode of failure in the contact zone that requires a consistent and correct definition of compressive material properties, both in the elastic range and in the post-failure condition. The data introduced in the simulation were instead extrapolated by a generic database, so a certain dispersion in results was expected when comparing the experimental failure load of about 40 kN to the numerical one, as in Table 9. The post-failure parameters (ERODS and SLIM\({}_{xx}\)) characteristic of MAT58 could also play a role in the correct identification of the failure mode; a correct calibration seems to be required on the basis of the real characterization of the material properties. An extensive characterization of the material properties seems necessary with a deeper calibration of the principal post-failure numerical parameters available in the proposed LS-DYNA material models. The adoption of the same erosion parameters for the two situations could also be too stringent assumption and a more detailed investigation on its sensitivity to the numerical simulation will need to be performed as a subject for future research activities. The punctual strain values obtained in the proposed numerical model at 20 kN of tensile load for both composite specimens were reported in [PERSON] et al. [12] against the results proposed by [PERSON] et al. [1], as summarized in Figures 12 and 13. The deformation local path is in good agreement with the literature results for SG 3-4 and SG 5. The strain field differences between the advanced and simple composite joints have also been found for numerical SG 1-2. However, a slight difference between the results found by [PERSON] et al. and the present value was observed. The lack of a complete material characterization was the major effect that could influence the obtained numerical results. The global strain field maps between the advanced and simple composite joints are compared in Figure 14. The picture reports the localized longitudinal strain map obtained at 20 kN of tensile load in the external composite ply of the proposed joints. A certain level of strain-concentration is evident near the contact area and at the hole edges in the simple configuration (B). A large reduction in the hole edge concentration and homogeneous distribution in the entire external surface can be evaluated with the application of the advanced solution (A). Reduced differences have been found in comparison to the literature results. A qualitative view of the longitudinal strain distribution was proposed. An extended evaluation of stress-strain is proposed for future work. Figure 12: Strain comparison for SG 1–2 and 3–4 between static FE [2] and present numerical simulation. Experimental strain values are from [PERSON] et al. [1].