## 6 Experimental Results and Discussion
### Hysteretic Response
& \multicolumn{1}{c}{Load} & \multicolumn{1}{c}{Max} & \(\mathbf{V_{acc}}\) & \(\mathbf{V_{acc}}\) & \(\mathbf{V_{acc}}\)  
In Figures 13 and 14, positive values correspond to the downward-loading direction (i.e., slab in tension), while negative values correspond to upward-loading (i.e., slab in compression). The experimental load-deformation responses are compared with the analytical shear and moment capacity estimates presented in Section 4. Figure 14 also provides the generalized backbone curve per Table 10-7 in ASCE 41-13, which accounts for cyclic degradation, for comparison with the experimental cyclic backbone. The plotted ASCE 41 backbones (in blue) assume a flexure-controlled performance, which is appropriate for all three FRP-strengthened specimens and the following modeling parameters: a = 0.025 rad; b = 0.05 rad; c = 0.20). The yield point was determined as the intersection of a line with slope m = M/\(\Theta\) = (2/Ls)*Elfhex, and the effective yield moment obtained from moment-curvature analysis is provided in Table 5. Hereby, M is the moment, \(\Theta\) is the chord rotation (i.e., equivalent to drift), and Elfhex is the prescriptive effective flexural stiffness for RC beams per Table 10-5 in ASCE 41-13 (i.e., Elfhex = 0.30E\({}_{\text{c}}\)I\({}_{\text{g}}\)). A 5% hardening slope was assumed to obtain the peak point. The acceptance criteria for the tested beams are also defined in Table 10-7 in ASCE 41-13 and consist of plastic rotation limits of 0.025 rad (i.e., 2.5.% drift) for the Life Safety (LS) performance objective (i.e., peak  
Figure 13: Shear-displacement hysteretic responses vs. shear capacity per ACI 318 and ACI 440.2R: (**a**) specimen S1A; (**b**) specimen S1B; (**c**) specimen S5. point of the ASCE 41 backbone) and 0.05 rad (i.e., 5.0% drift) for the Collapse Prevention (CP) performance objective (i.e., point of ultimate rotation and residual strength). All three specimens showed a more flexible elastic response than predicted per the ASCE 41 backbone (Figure 14). This is mainly due to the fact that the analytical elastic stiffness used to construct the ASCE 41 backbone curve (i.e., \(\text{EI}_{\text{flex}}=0.30\text{E}_{\text{c}}\text{I}_{\text{g}}\)) only accounts for flexural deformations. As discussed in Section 6.5, the contribution of shear deformations was derived from the LVDT readings and was found to be significant for all three specimens (up to 40% of the total tip deflection). None of the specimens reached the analytically predicted flexural capacity (\(\text{M}_{\text{u}}\); Section 4.1), which may be partially explained by the fact that the \(\text{M}_{\text{u}}\) values correspond to monotonic loading; thus, they do not account for cyclic strength degradation. For specimens S1A and S1B, the deficit of the experimental ultimate moments with respect to the analytical predictions is three times larger in the downward-loading direction (for S1A, 12% down and 4% up; for S1B, 11% down and 4% up), which the authors attributed to the inefficacy of either of the U-wrap schemes in providing any shear strengthening to the beams in the downward-loading direction. Specim S5 achieved a flexural capacity that was just 5% below the expected  
Figure 14: Moment-rotation hysteretic response vs. ASCE 41 backbone curves: (a) specimen S1A: (b) specimen S1B; (c) specimen S5. (monotonic) ultimate moment in the downward direction, but twice as much, 11%, in the upward direction. Furthermore, the proposed full-wrap EB-FRP-reinforcing scheme tested on specimen S5 enabled the development of considerably high rotation values--0.034 rad (3.4% drift) at peak and 0.038 rad (3.8% drift) at the onset of strength loss (due to FRP rupture) for an MRF beam that, otherwise, would have prematurely failed in shear (due to 30% shear overstress in the downward-loading direction, Figure 13c).