Objective: Curved aramid fiber-reinforced composites are important load-bearing and energy-dissipating components in protective helmets and other shell-type protective structures. Unlike flat laminates, curved shells resist transverse impact through the coupled actions of local indentation, bending, membrane stretching, and geometric flattening. Therefore, impact indices obtained from flat panels cannot fully characterize the load transfer and damage evolution of a locally spherical shell. Specifically, the coupled effects of bending radius (R), ply count (L), and impact energy (E) on the transition from localized failure to global deformation remain insufficiently understood. This study aimed to clarify how these parameters govern the low-velocity impact response of curved aramid composites, establish a correspondence between response curve signatures and physical damage, and provide an experimental basis for the structural design of curved personal protective equipment. Methods: Curved aramid/epoxy laminates were fabricated from 1500D aramid prepreg on custom split stainless-steel spherical molds by vacuum bag/autoclave molding. The 200 mm × 200 mm prepreg sheets were laid onto the molds ply by ply, locally softened during forming to reduce wrinkling and trapped air, vacuum-sealed, and consolidated at a pressure ramp of 0.09 MPa/min to a maximum pressure of 3 MPa. A full test matrix was constructed using bending radii (R) of 80, 120, and 160 mm, ply counts (L) of 6, 8, and 10, and nominal E values of 12, 17, and 22 J. Drop-weight impact tests were performed according to ASTM D7136/D7136M using a 5.5 kg impactor fitted with a 16 mm diameter hemispherical tup; the curved specimens were restrained at four corners. Force–displacement, force–time, and energy–time histories were recorded. Maximum dynamic displacement (U_max), peak impact force (F_max), time to peak force, and energy absorption ratio were extracted and compared across the parameter combinations. Post-impact damage morphologies were then correlated with the stages identified from the response curves. Results: Two reproducible response modes were identified. Mode Ⅰ, which occurred mainly in specimens with a larger radius or fewer plies, exhibited an initial slow increase and slight decrease, followed by a rapid increase to a relatively high peak and an abrupt post-peak drop. Its deformation was initially dominated by local indentation and matrix cracking; subsequent yarn stretching and breakage induced concentrated delamination and sudden loss of load-carrying capacity. Mode Ⅱ, favored by a smaller radius or a larger ply count, showed a gradual increase, a stable plateau, and a slow unloading stage. Wide-area bending, yarn rotation and stretching, interfacial sliding, and crack deflection across multiple interfaces enabled a greater proportion of the specimen to participate in load bearing, producing distributed and progressive damage. At fixed values of E and L, increasing R reduced U_max but increased F_max. For the six-ply specimens, increasing R from 80 to 120 mm reduced U_max by 18.07%–21.77% across the three impact energies, whereas increasing R from 80 to 160 mm reduced U_max by 33.26%–34.98%. Under the 12 J, six-ply condition, the time to F_max decreased from 33.82 ms at R = 80 mm to 23.76 ms at R = 120 mm and 17.66 ms at R = 160 mm. Thus, the flatter specimens developed a higher peak force more rapidly and with less global displacement but were more susceptible to abrupt local failure. The sensitivity to E also depended on the response mode. Relative to 12 J, U_max of the R80-L8 specimens increased by 37.57% at 17 J and by 70.38% at 22 J, while the corresponding increases for R80-L10 were 18.80% and 65.07%. When E was increased from 12 to 22 J, F_max increased by 13.25%–64.03% for Mode Ⅰ configurations but remained within 743.79–776.59 N for Mode Ⅱ R80-L8 specimens, indicating a comparatively stable force response. The dissipated energy fraction generally increased with E, and at 22 J, it was approximately 90% or higher for most specimens. Increasing L, however, tended to reduce this fraction, indicating that a stiffer laminate retained a greater recoverable elastic energy share. For example, at R = 80 mm and E = 12 J, the energy absorption ratio decreased from 99.0% for six plies to 82.5% for eight plies, with a concurrent transition to Mode Ⅱ. The damage observations confirmed that a high dissipated energy ratio should be interpreted together with the spatial concentration and severity of irreversible damage. Conclusions: Bending radius and ply count jointly determine whether a curved aramid laminate resists low-velocity impact through localized high-load bearing or through global, progressive deformation. A larger radius reduces maximum displacement and increases peak force, but it also aggravates contact zone deterioration and promotes abrupt failure. A smaller radius facilitates load redistribution and global bending, although it allows a larger overall deflection. Additional plies improve cooperative load sharing and stabilize progressive damage, while the associated increase in stiffness can lower the dissipated energy fraction. Consequently, curved protective structures should not be optimized solely by peak force, displacement, or absorbed energy. Curvature and laminate thickness must be coordinated to balance force transmission, allowable deformation, and damage localization. The proposed two-mode classification connects measurable response curves with damage mechanisms and provides a practical framework for optimizing aramid composite helmet shells and related curved protective components.