## 4 Discussion  
The experimental information presented in Figures 10 and 11 is here discussed, to understand the stress sharing of the masonry-frame system, so as to highlight the main concerns regarding the design procedure of such retrofit solutions. To this purpose, the test on R1 in the retrofitted system (which got the clearest graphs of the two tests) is compared to the test on its unretrofitted behavior. Figure 14 reports in blue the values of tense diagonal displacement (Figure 10) and, in magenta, the drift (Figure 6) normalized to the respective maximum values of unretrofitted masonry (N1). The plot of displacements of N1 along the compressive diagonal was too disturbed (Figure 10) to provide clear values and thus the plot of the compressed diagonal was not selected for this comparison. The histogram clarifies the gain of overall stiffness and displacement capacity that this kind of retrofit could bring on. The solid blue columns show that R1 allows for higher displacements along the tense diagonal of the wallette without cracking; in fact, R1 shows a +71% higher displacement at 5.28 kN (i.e., the load of first crack for N1), and a +27% higher displacement at 8.59 kN (peak of N1) without cracking in the masonry. After cracking, at 11 kN, the diagonal displacement increases by up to 4.8 times the reference value. The magenta columns allow us to appreciate the great difference in drift up to 8.59 kN (\(-\)88% for the retrofitted wallette), and beyond, up to the abrupt failure at 16.69 kN.  
The histogram in Figure 15 is construed after calculating the axial load from the strain gauge plots of the GFRP profiles, assuming the elastic modulus value in Table 2. The tensile load (vertical profile) is normalized by the maximum load \(N_{max}\) (Equation (8)), and the compressive load (diagonal profile) by the critical load \(N_{cr}\) (Equation (9)) for the specific length of the profile segment:  
\[N_{max}=\frac{f_{c,long}\cdot A}{\gamma_{m,f}}, \tag{8}\]  
\[N_{cr}=\frac{N_{max}}{1+\frac{N_{max}}{N_{Bol}}} \tag{9}\]  
where \(f_{c,long}\) is the profile's longitudinal compressive strength, \(\gamma_{m,f}\) the safety factor (assumed to be equal to 1.3), \(A\) the cross-sectional area and \(N_{Eul}\) the Eulerian load [35].  
Figure 13: Updated FE model of the case study; crack pattern compared to the test results.  
Despite the noticeable increase in strength provided by the frame retrofit, Figure 13 points out the very low stress rates of the GFRP profiles, slowly increasing up to failure and not exceeding 2% of the profile's maximum capacity. The low exploitation of the GFRP's strength was expected, since the masonry remains the weakest ring even in the retrofitted system; however, the stress rate on the retrofit frame surely deserves further investigation, possibly by numerical means, to understand the facts of stress transfer from the masonry to the frame through the anchors (which it was not possible to monitor during the tests). The frame is thus very effective in impairing the weakest mechanism of failure, mainly due to the vertical profile; the diagonal member of the frame shows a weaker relief of the stress, as one can also see from Figure 11 (left). Indeed, the choice of a cross-section of much higher  
Figure 14: Comparison between the displacements of unretrofitted and retrofitted masonry across tests N1 and R1.  
Figure 15: Ratio of axial load to maximum load (vertical profile) and critical load (diagonal profile) in the GFRP frame at different load levels.  
capacity than the design buckling load--mainly due to the practical feasibility of the connections--has led to a very stiff frame and a low rate of work in the pultruded profiles.