## 3 Results  
The force-drift graphs of the three tests N1 (unretrofitted), R1 and R2 (retrofitted) are plotted in Figure 6. The force and displacement values at the end of the first crack, peak and test are detailed in Table 4.  
In Table 4, the predictions of the FE models are reported under the respective cases of the masonry and the masonry-frame system. It can be seen that Test N1 gave a pretty good confirmation of the analytical and FE predictions, as far as they concerned the maximum load. Because of the constraints of the physical model, the corner uplifting at the second bed joint from the bottom, at 8.59 kN, defined the shear strength \(V_{r}\) (Figure 7). The wallette, being very slender, collapsed shortly after. Tests R1 and R2 were performed on equal physical models of the retrofitted wallette. The longitudinal strain of the FRP profiles was monitored with PL-60-11 strain gauges. Tests R1 and R2 gave very similar results in term of failure mode and \(P_{max}\) (respectively, 16.62 and 15.9 kN); with respect to test N1, the retrofitted wallettes revealed a remarkably different behavior and dramatic increase in the ultimate horizontal load. In detail, for both R1 and R2, the cracking initiation was at the half-height of the wallette; the sawtooth crack propagated towards the bottom along an inclined direction and stopped at one third of the wallette's width. Although this does not fully match any of the usual modes, it is clear that the retrofitted wallettes do not encounter a rocking mechanism. As for the initial stiffness, the retrofitted condition shows a small but appreciable increase with respect to the unretrofitted wallette. On the other hand, the load-drift paths of R1 and R2 show some differences between each other. In fact, R1 collapsed under testing (Figure 8), while R2 reached a dramatically higher ultimate displacement and finally survived the test (Figure 9). The lesser drift performance of \begin{table}
\begin{tabular}{c c c c c c c} \hline \hline \multirow{2}{*}{**Case**} & \multicolumn{2}{c}{**First Crack**} & \multicolumn{2}{c}{**Maximum Force**} & \multicolumn{2}{c}{**Ultimate Displacement**} \\ \cline{2-7}  & \(P_{1}\) **(kN)** & \(d_{1}\) **(mm)** & \(P_{max}\) **(kN)** & \(d_{2}\) **(mm)** & \(P_{u}\) **(kN)** & \(d_{u}\) **(mm)** \\ \hline N1 & 5.275 & 1.401 & 8.594 & 13.165 & 8.081 & 14.287 \\ N-FE & 8.000 & 0.106 & 8.000 & 0.106 & 9.000 & 0.452 \\ R1 & 11.899 & 3.034 & 16.621 & 30.277 & 16.621 & 30.277 \\  & (+125\%) & (+117\%) & (+93\%) & (+130\%) & (+106\%) & (+112\%) \\ R2 & 8.068 & 1.951 & 15.904 & 21.818 & 16.010 & 57.603 \\  & (+53\%) & (+39\%) & (+85\%) & (+66\%) & (+98\%) & (+303\%) \\ R-FE & 6.000 & 0.060 & 16.000 & 0.315 & 17.000 & 0.485 \\ \hline \hline \end{tabular}
\end{table} 
Table 4: Results of the tests and per cent variation of retrofitted versus unretrofitted case. Figure 6: Horizontal force–drift graph of the three tests. R1 is likely due to the geometric imperfections of the wallette, which was found to be visibly out of plumb after seasoning. As with the unretrofitted case, the experimental results are in a rather good agreement with the predictions of the FE analysis as far as they concern the maximum shear force of both the masonry and the masonry-frame system; in particular, the strength increase due to the retrofit (calculated on the basis of the average of \(P_{max}\) of R1 and R2) is +89%, which equals the FE prediction. On the other hand, the FE models could not catch the experimental shear stiffness and displacements magnitude. This is due to the model input, consisting of tensile and compressive behavior parameters, which did not allow accounting for the shear stress-drift relation explicitly. The recordings of transducers along the wallette's diagonals revealed the effectiveness of the FRP frame retrofit in retaining the displacements along the tense diagonal (Figure 10).