## 2 Materials and Methods
\begin{table}
\begin{tabular}{c c c c} \hline \hline
**Material** & **Property** & **Value** & **Source** \\ \hline \multirow{2}{*}{Brick} & Compressive strength \(f_{\text{dc}}\) (N/mm\({}^{2}\)) & 20.00 & Destructive testing \\  & Elastic modulus \(E_{b}\) (N/mm\({}^{2}\)) & 5700 & Destructive testing \\  & Poisson’s ratio \(\
u_{b}\) (-) & 0.18 & producer’s data \\ Mortar & Compressive strength \(f_{\text{mc}}\) (N/mm\({}^{2}\)) & 4.40 & Destructive testing \\  & Flexural strength \(f_{\text{mf}}\) (N/mm\({}^{2}\)) & 0.92 & Destructive testing \\  & Elastic modulus \(E_{\text{m}}\) (N/mm\({}^{2}\)) & 5690 & Destructive testing \\  & Poisson’s ratio \(\
u_{m}\) (-) & 0.25 & producer’s data \\ Masonry & Compressive strength \(f_{\text{mc}}\) (N/mm\({}^{2}\)) & 14.96 & Equation (3) \\  & Elastic modulus \(E_{\text{up}}\) (N/mm\({}^{2}\)) & 4878 & Equation (4) \\  & Joint shear strength at \(\sigma_{0}\) = 0.1 N/mm\({}^{2}\)\(\tau_{0.1}\) (N/mm\({}^{2}\)) & 0.56 & Destructive testing \\  & Joint peak shear displacement \(d_{\text{peak}}\) (mm) & 1.6 & Destructive testing \\ \hline \hline \end{tabular}
\end{table} 
Table 3: Material properties. Figure 3: Transducers and gauges locations and physical models N1, R1, and R2 under testing setup. In Table 3, the masonry compressive strength \(f_{\text{zuc}}\) was evaluated according to [PERSON]'s theory [32]:  
\[f_{\text{zuc}}=\frac{f_{\text{zuc}}}{1+\frac{a(\eta_{\text{z}}-b\eta_{\text{zuc}} )}{k(1+ab-\
u_{\text{z}}-ab\eta_{\text{z}})}}, \tag{3}\]  
where \(f_{\text{zuc}}\) is the compressive strength of the brick, \(a\) is the ratio between the thickness of bed joint and brick (\(t_{\text{m}}/t_{\text{b}}\)), \(b\) the ratio of the mortar's elastic modulus to the brick's elastic modulus (\(E_{\text{m}}/E_{\text{b}}\)), and \(k\) is the ratio between the brick's tensile and compressive strength (\(f_{\text{zuc}}/f_{\text{b}t}\) assumed equal to 0.1). It turns out to be \(f_{\text{zuc}}=14.96\) N/mm\({}^{2}\) as reported in Table 2.  
The elastic modulus of masonry \(E_{\text{zuc}}\) along the direction of axial compression is calculated with the expression reported by [PERSON] et al. [33]  
\[\frac{1}{E_{\text{zuc}}}=\frac{\eta_{b}}{E_{b}}+\frac{\eta_{m}}{E_{m}}, \tag{4}\]  
where \(\eta_{b}\) and \(\eta_{m}\) are the per cent fraction of total unit plus the joint thickness of each component. The collected data show a masonry of average to good quality in compression, while the shear behavior may be poorer because of the low tensile properties of the mortar. Before the beginning of the tests, a prediction of the maximum shear load had been made. Considering the square shape of the panel, the following expressions for rocking failure load (\(V_{r}\)) and toe crushing (\(V_{\text{tc}}\)), proposed by Tassios [34], were applied:  
\[V_{r}=\frac{l^{2}t}{2h}\bigg{(}\sigma_{0}-\frac{\sigma_{0}}{f_{\text{zuc}}} \bigg{)}, \tag{5}\]  
\[V_{\text{tc}}=\frac{f_{\text{zuc}}-\sigma_{0}}{6h}l^{2}t, \tag{6}\]  
where \(l\), \(t\) and \(h\) are the wallette's dimensions as in Figure 2B, and \(\sigma_{0}\) is the vertical compressive stress due to gravity and superimposed load. The minimum of the calculated values, i.e., \(V_{r}=9.056\) kN, indicates a rocking failure. Furthermore, an FE model was set up to get a picture of the expected structural behavior of the physical model, and to evaluate the potential performance improvement attainable with the retrofit. The structural analysis was performed with DIANA FEA Finite Element code, version 9.6. According to the small size of the physical model, the bricks and joints are reproduced in detail, to account for the orthotropy of the masonry and to catch the details of the crack pattern. Quadilateral eight-node shells simulated the masonry components, being suitable for nonlinear analysis; two-node beam elements simulated the FRP profiles and connection bars.