To support the seismic optimization of long-span bridges in regions of high seismicity, this study evaluates the seismic performance of continuous rigid-frame box-girder bridges with different web configurations. A continuous box-girder bridge with corrugated steel webs (CSWBGB) having a main span of 105 m was analyzed and compared with two control models: a continuous box-girder bridge with flat steel webs (FSWBGB) and a conventional prestressed concrete box-girder bridge (PCBGB). Finite element models of the three web types were developed using MIDAS/Civil, and seismic responses were evaluated using the response spectrum method with geometric nonlinearity incorporated; the analyses were conducted under E1 and E2 ground motion intensities (corresponding to a 63% probability of exceedance in 100 years and a 2% probability in 50 years, respectively, as specified in the Chinese seismic design code). Displacement, axial force, and shear force responses were systematically compared among the three configurations. The results show markedly different seismic responses despite the bridges having similar fundamental frequencies. In the longitudinal direction under seismic excitation, the CSWBGB exhibited larger axial displacement than the FSWBGB, yet its peak axial force and shear force decreased by 13% and 18%, respectively, indicating that the greater axial deformation helps relieve internal force demands. Under transverse E1 seismic action, the CSWBGB displayed smaller lateral displacements than both the FSWBGB and the PCBGB. Compared with the CSWBGB, the PCBGB experienced an 11% larger longitudinal displacement and a 43% higher peak axial force, reflecting its relatively limited seismic performance. These findings demonstrate that the CSWBGB not only provides lighter self-weight than the PCBGB but also offers enhanced transverse stiffness, which results in smaller lateral displacements and lower peak shear forces—thus achieving an optimal balance between lightweight design and structural strength. Although the CSWBGB shows strong potential for practical application, its longitudinal displacement response should be carefully controlled in design. 1. Introduction In recent years, corrugated steel webs (CSWs), as an innovative structural form that combines lightweight characteristics with high shear strength, have offered a potential technical solution for enhancing the seismic performance of bridges. By replacing concrete webs with CSWs, the structural self-weight can be reduced while fully exploiting the compressive strength of concrete and the tensile strength of steel, resulting in advantages such as light weight, high strength, and ease of construction [ 16, 17, 18, 19, 20]. Compared with the traditional FSWBGB, the continuous box-girder bridge with corrugated steel webs (CSWBGB) exhibits significantly improved out-of-plane stiffness of the web section due to its unique corrugated configuration. This not only effectively suppresses web buckling deformation, thereby enhancing overall structural stability, but also increases the shear capacity of the web through the “accordion effect” [ 21]. The accordion effect, arising from the axial flexibility of the corrugated geometry, decouples the web’s longitudinal deformation from its shear resistance, enabling efficient shear force transfer while minimizing bending stiffness participation. Lindner and Aschinger [ 22] noted that trapezoidal webs could substantially increase the initial shear buckling stress compared with FSWs. Bi et al. [ 23] demonstrated experimentally that CSWs can effectively avoid the safety risks associated with FSWs, which are prone to buckling failure without obvious warning before reaching their ultimate load capacity. Zha et al. [ 24] further pointed out that in CSW composite beams, the corrugated geometry yields a more uniform shear stress distribution along the web height, effectively avoiding the stress concentration commonly observed in traditional flat webs, and that the corrugated configuration provides excellent buckling stability, significantly reducing the risk of local buckling. Extensive research has been conducted on the mechanical behavior of CSWBGBs. Regarding shear performance, Peng et al. [ 25] indicated that CSWs provide higher out-of-plane stiffness and shear buckling strength than FSWs, thereby enhancing web stability. Bi et al. [ 26] systematically elucidated the accordion effect, finding that under combined bending and shear, the flexural and shear capacities can simultaneously reach their maxima without significant interactive reduction. Li et al. [ 27] revealed the mechanical behavior of a prefabricated CSWBGB under static loading through experimental and theoretical investigations, demonstrating that CSWs optimize shear stress distribution and significantly contribute to improved shear efficiency and self-weight reduction. In terms of axial–flexural behavior, Huang et al. [ 28] pointed out that CSWs readily undergo compressive or tensile deformation under axial forces, but their longitudinal bending stiffness is extremely low, so the longitudinal bending moment they carry is negligible. This characteristic allows prestress to be more efficiently introduced into the concrete top and bottom slabs, greatly improving material utilization efficiency. Wang et al. [ 29] proposed a method considering the accordion effect for calculating the web shear stress in composite box beams with variable cross-sections and noted that the influence of the shear-lag effect cannot be neglected [ 30, 31]. For seismic analysis, Huang et al. [ 32] employed time-history analysis to evaluate the seismic performance of a long-span continuous rigid-frame bridge with CSWs. Currently, a growing number of researchers have employed machine learning methods to evaluate the seismic performance and damage effects of frame structures under earthquake action, providing new technical approaches and tools for subsequent seismic research. However, these methods have not yet been applied to bridge structural systems, and their applicability to different bridge types remains to be further investigated [ 33, 34]. Moreover, simplified analytical approaches have been developed, including equivalent calculation models and deflection calculation methods for CSWBGBs proposed by Feng et al. [ 35], and a forward inclination angle calculation method for CSWBGBs by Wang et al. [ 36] that accounts for both shear-induced deflection and additional shear deflection caused by bending moment. Despite this extensive body of research, most studies have focused on static loading conditions or the behavior of individual components. Systematic comparative investigations into the seismic performance of continuous rigid-frame bridges with different web types (CSWs, FSWs, and PC webs) at the full bridge system level remain relatively scarce [ 37, 38]. The influence of web type on the overall displacement response, internal force distribution, stress state, and damage mechanisms under rare earthquake events has not been fully elucidated. This gap hinders an accurate assessment of the comprehensive advantages offered by CSWBGBs in seismic regions. Therefore, this study takes a CSWBGB with a main span of 105 m as the research object and establishes corresponding finite element (FE) models of bridges with FSWs and PC webs. The response spectrum method is employed to systematically compare and analyze the seismic performance of the three web configurations under E1 and E2 seismic actions. The influence mechanism of web type on the seismic response of long-span continuous rigid-frame bridges is investigated, aiming to provide theoretical basis and data support for web type selection and seismic optimization design of such bridges in high-intensity seismic zones. 2. Practical Bridge Project The case is a four-span continuous box girder bridge located in Beijing, China. It adopts a CSW–prestressed concrete composite box girder structure (CSWBGB) and has been completed. The total length of the bridge is 330 m, with a span arrangement of 60 + 2 × 105 + 60 m, consisting of two side spans of 60 m and two main spans of 105 m, as shown in Figure 1. The superstructure is a prestressed concrete continuous box girder in which CSWs serve as the primary shear-resisting components. The main girder features a single-cell single-box cross section with straight webs. The top slab has a total width of 13.75 m, with cantilevers of 3.25 m on each side. The girder depth is 7.0 m at the intermediate supports and 3.5 m at the side supports and in the mid-span regions, varying longitudinally according to a 1.8-power parabolic curve. The top slab thickness is 0.32 m, while the cantilever slab thickness varies from 0.20 m at the tip to 0.80 m at the root following a polygonal broken line. The bottom slab thickness is 0.32 m at mid-span and 0.80 m at the supports, also following a 1.8-power parabolic curve. Owing to the complex stress state in segment #0, and to ensure effective force transfer, the concrete web thickness in this segment is set to 1.2 m, the top slab thickness to 1.9 m, and the bottom slab thickness to 1.4 m; fillets are provided at the web–bottom slab connections. For construction, the main girder is divided longitudinally into 11 segments. Excluding the pier-top segment above the central pier (Seg. #0), the side-span support segment (Seg. #11), and the closure segment (Seg. #10), segments #2 to #9 are built using the balanced cantilever cast-in-place method with formwork travelers. These cantilever segments consist of five 4.8 m-long segments and three 6.4 m-long segments. Compared with the traditional cantilever casting method, the formwork travelers are supported directly on the steel girder, and the standard segment length is increased from 4.8 m to 6.4 m, effectively shortening the construction period. The box girder is provided with three diaphragms in each side span and six diaphragms in each main span, all 0.50 m thick and longitudinally spaced at distances ranging from 12.8 m to 16.7 m. In addition, to improve local mechanical performance, lining concrete is cast inside the CSWs in segments #1 and #2 on both sides of the central pier and in segment #11 at the transition pier. Details of the connection between the CSW and the upper and lower concrete slabs are shown in Figure 2. The connection between the CSW and the top slab concrete is achieved using a double Twin-PBL connector. The thickness of the CSWs varies from 24 mm at the pier to 20 mm at mid-span. The CSWs have a wavelength of 1.60 m, a wave height of 0.22 m, and a horizontal panel width of 0.43 m. The horizontal fold angle is 30.7°, and the bend radius is 15t, where t is the CSW thickness. The typical box-girder cross section and the standard corrugation profile are shown in Figure 3a. To investigate the influence of web type on the seismic response of the bridge, the CSWs of the CSWBGB are replaced with flat steel webs (FSWs) and prestressed concrete webs, respectively, while keeping all other structural parameters (such as dimensions and material properties) unchanged. The cross section of the resulting FSWBGB is shown in Figure 3b. For the PCBGB, the economical web thicknesses commonly adopted for traditional concrete bridges of comparable spans were consulted (see Table 1). Based on the span arrangement of the studied bridge (60 m + 2 × 105 m + 60 m), a web thickness of 450 mm was selected. The corresponding cross section is shown in Figure 3c. 3. FE Modeling 3.1. Description of Finite Element Model To investigate the seismic performance of continuous box girder bridges with different web configurations, a full bridge finite element model was developed using MIDAS/Civil 2020 [ 39]. The following simplifications are adopted in the modeling: (1) the connection between the CSW and the concrete top and bottom slabs is fully reliable, with no relative slip or shear failure; (2) the CSW possesses sufficient strength to preclude shear buckling or lateral instability; (3) the bonded prestressing tendons are perfectly bonded to the concrete, while the external prestressing tendons are effectively connected to the structure at anchor points and deviators without slip; (4) prestress losses are temporarily neglected. Both the box girder and the piers are modeled using beam elements. To account for time-dependent effects of concrete, the basic creep and shrinkage prediction model recommended by Eurocode 2 is employed [ 40]. It is noted that MIDAS/Civil provides a dedicated beam element that reasonably simulates the mechanical contribution of the CSW through an equivalent stiffness approach [ 41]. This element is also applicable to the simulation of prestressed concrete girders with CSWs and variable cross-sections. The full-bridge model consists of 547 nodes and 558 elements (see Figure 4). The main girder adopts a variable-depth box girder, with the girder depth varying from 7.0 m at the supports to 3.5 m at the mid-span following a 1.8-power parabolic curve. This variable cross-section is achieved using the ‘variable cross-section group’ function in MIDAS/Civil: section parameters are defined at each node, and each construction segment (with a length of 4.8 m or 6.4 m) is modeled as a single beam element, with different sections at the two ends of the element to accurately reflect the continuous variation of girder depth. The variable cross-sections of the piers are treated in the same manner. The boundary conditions are set as follows: fixed constraints at the pier bottoms, simulating the connection to the pile foundations; and roller supports at both ends of the side spans. The longitudinal restraint arrangement of the piers is explicitly considered in the finite element model. In the finite element model, the left-side superstructure is rigidly connected to the pier column, while the right-side pier and superstructure are connected via movable supports. This arrangement is consistent with the actual restraint system of the bridge. Consequently, under longitudinal seismic excitation, the inertial force of the superstructure is mainly transferred to the fixed-pier system, while the movable pier releases part of the longitudinal constraint. Material properties are assigned according to the design specifications, and the key parameters (including elastic modulus, Poisson’s ratio, coefficient of thermal expansion, and unit weight) are listed in Table 2. All steel components are Q345 steel, and the concrete components are C55 concrete. To account for the interaction between axial compressive forces and lateral displacements in this pier–girder rigid frame system, geometric nonlinearity was activated in the MIDAS/Civil model, with the geometric stiffness matrix incorporated into both the eigenvalue analysis and the subsequent response spectrum analysis. Given that the superstructure concrete slabs are subjected to considerable tensile stress demands under seismic loading, particularly under E2 action, the structural response of the superstructure may extend into an elasto-plastic range; accordingly, the seismic response results presented herein are interpreted within this context, with the comparative trends among the three web configurations remaining the primary focus of the analysis. Furthermore, a detailed construction stage analysis was conducted in this study. Following the actual construction sequence, structural elements were activated stage by stage with corresponding loads applied, including structural self-weight, secondary dead load (deck pavement, railings, etc.), prestressing tension forces, and construction formwork traveler loads. The tensioning force of each prestressing tendon is obtained by multiplying the tensile stress by its cross-sectional area. The initial internal forces and initial displacements in the completed bridge state were obtained through the construction stage analysis. Subsequently, this completed bridge state was used as the initial condition in the response spectrum analysis, with the seismic action superimposed upon it. This approach ensures that the initial internal forces and deformations induced by dead load and construction loads work simultaneously with the seismic action, faithfully reflecting the actual stress state of the structure. Three bridge configurations including CSWBGB, FSWBGB, and PCBGB were established in this study (see Figure 5). For each configuration, the seismic response under the two fortification intensity levels, E1 and E2, was calculated using the corresponding response spectra, and a systematic comparative analysis was conducted. The response spectrum method was adopted in accordance with the Chinese Specifications for Seismic Design of Highway Bridges [ 42], which prescribes this approach as the standard procedure for Seismic Fortification Category B bridges with regular span configurations such as those examined here; it also ensures a consistent and directly comparable analytical basis across all three web configurations. 3.2. Structural Seismic Characteristic Parameters According to the Chinese Specifications for Seismic Design of Highway Bridges [ 42], the bridge, with a single span not exceeding 150 m, is classified as Seismic Fortification Category B. Accordingly, a two-level fortification and two-stage design approach is adopted. The seismic design of this bridge is based on an exceedance probability of 63% in 100 years for E1 and 2% in 50 years for E2. The peak ground acceleration is 0.10 g, the characteristic period of the basic seismic acceleration response spectrum is 0.45 s, and the basic seismic intensity is VII degrees. For engineering seismic design purposes, the response spectrum obtained from the site-specific dynamic response analysis of the soil layers is normalized. The resulting design response spectrum is given by Equation (1). S a ( T ) = A max · β ( T ) (1) where Amax is the peak ground acceleration of the design ground motion and β( T) is the design acceleration amplification factor defined by Equation (2). β ( T ) = 1 T ≤ 0.04 s 1 + ( β m − 1 ) T − 0.04 T 1 − 0.04 0.04 s ≤ T ≤ T 1 β m T 1 ≤ T ≤ T g β m ( T g − T ) γ T g ≤ T ≤ 6 s (2) where, Tg is the characteristic period, βm is the amplification coefficient, γ is the decay coefficient, and T1 is the period at which the constant-acceleration segment ends. The design ground-motion parameters corresponding to different exceedance probabilities for the site are listed in Table 3. The acceleration response spectrum Sa( T) for the E1 seismic action is shown in Figure 6a, with a maximum spectral acceleration Smax = 0.1709 g. The spectrum for the E2 seismic action is shown in Figure 6b, with Smax = 0.651 g. These acceleration response spectra were applied as seismic input in the X, Y, and Z directions of the bridge model. 4. Modal Analysis Modal analysis was performed as an indispensable prerequisite for the subsequent response spectrum analysis: the natural frequencies, mode shapes, and modal mass participation ratios extracted herein serve as direct inputs for computing the seismic response. The Ritz vector method was employed, as it is computationally efficient and particularly well-suited for response spectrum analysis applications. Modal analysis was performed to determine the natural frequencies and mode shapes required for the seismic response spectrum analysis. The first three modes of the CSWBGB were selected for comparison with those of the FSWBGB and PCBGB. In the response spectrum analysis, a total of 30 Ritz vectors were extracted for each bridge configuration to ensure sufficient modal contribution in the principal directions. The cumulative modal mass participation ratios of the three bridge models are summarized in Table 4. As shown in the table, the cumulative mass participation ratios in the longitudinal, transverse, and vertical directions all exceed 90%, satisfying the commonly adopted requirement in seismic design codes. Therefore, the extracted modes are considered sufficient for accurately representing the dynamic response of the bridge system in the subsequent response spectrum analysis. For the three box-girder bridges with different web configurations, the dynamic characteristics depend primarily on the web type and the resulting sectional stiffness, which jointly determine the natural frequencies, mode shapes, and dynamic response characteristics. As shown in Table 5, the natural periods and frequencies of the three bridge types are similar. The maximum frequency difference between the CSWBGB and FSWBGB is approximately 7%, and that between the CSWBGB and PCBGB is about 3%. Beyond the third mode, the natural frequency of the PCBGB slightly exceeds that of the CSWBGB, though the difference remains marginal (roughly 1.4%). 5. Analysis of Seismic Response Results 5.1. Main Girder Displacement To investigate the deformation patterns of the three bridge configurations under seismic excitation, representative cross sections were selected for analysis: the mid-span section, the support section, and the quarter-span section, as illustrated in Figure 10. Under E1 longitudinal seismic action, the deformation patterns of the three bridges are shown in Figure 11. Among the three bridges, the PCBGB exhibited the largest longitudinal displacement under E1 seismic action, reaching approximately 10.74 mm, followed by the CSWBGB at 10.33 mm and the FSWBGB at about 9.49 mm. The FSWBGB, owing to its high axial stiffness and strong resistance to axial deformation, displayed approximately 11% smaller longitudinal displacement than the PCBGB. In contrast, the CSWBGB, with its corrugated web geometry, possesses relatively lower axial stiffness, resulting in approximately 9% greater longitudinal displacement than the FSWBGB-indicating that the corrugated configuration reduces axial restraint and permits greater axial deformability. The transverse deformation under E1 action is shown in Figure 12. The corresponding lateral displacements were approximately 6.61 mm for the CSWBGB, 8.84 mm for the FSWBGB, and 7.75 mm for the PCBGB, with the FSWBGB exhibiting the largest deformation among the three. The FSW has relatively low out-of-plane stiffness and weak resistance to lateral deformation, leading to a maximum displacement approximately 14% larger than that of the PCBGB. By contrast, the folded geometry of the CSW acts as an effective stiffener, substantially enhancing the out-of-plane stiffness of the web and improving resistance to lateral deformation. As a result, the CSWBGB exhibited a maximum transverse displacement approximately 25% smaller than that of the FSWBGB, reflecting the marked transverse stiffening effect of the corrugated configuration. The vertical deformation under E1 action is shown in Figure 13. All three bridges exhibited vertical bending, with maximum deformation occurring at the mid-span of the deck slab. The peak displacements of the CSWBGB, FSWBGB, and PCBGB were approximately 2.98 mm, 1.99 mm, and 2.58 mm, respectively—the CSWBGB exhibiting the largest deformation. From a mechanical mechanism perspective: Due to the ‘accordion effect,’ the CSWBGB exhibits a significantly reduced vertical flexural stiffness, resulting in a larger displacement response under vertical seismic action. The FSWBGB, on the other hand, possesses a certain longitudinal flexural stiffness because of its flat steel webs, which assist the concrete top and bottom slabs in jointly resisting vertical bending; consequently, its vertical flexural stiffness is higher, and the displacement response is smaller. Although the PCBGB also has a relatively high vertical flexural stiffness, its concrete webs are thick, and its self-weight far exceeds that of the FSWBGB and CSWBGB; therefore, under the same vertical seismic input, it still undergoes a relatively large vertical displacement. Notably, the vertical displacements under seismic action were significantly smaller than the longitudinal and transverse displacements for all configurations, indicating that vertical response is not the governing factor in seismic design. Under E2 seismic action, the deformation characteristics of the three bridges were similar to those observed under E1. The maximum displacements of the PCBGB were 49.64 mm (longitudinal), 35.37 mm (transverse), and 10.88 mm (vertical). For the CSWBGB, the corresponding values were 47.11 mm, 30.37 mm, and 13.59 mm, while for the FSWBGB, they reached 41.70 mm, 40.32 mm, and 8.97 mm. These results confirm that, irrespective of web type, the dominant directions of the maximum displacement response remain longitudinal and transverse. From the perspective of engineering practice, it is therefore necessary to systematically enhance both the axial and lateral resistance of the bridge through coordinated overall stiffness design. For example, rationally arranged diaphragms in the mid-span region-through optimization of their spacing, cross-sectional form, and connection details with the main girder-can not only directly increase axial and lateral stiffness but also effectively restrain the distortion and warping effects of the main girder, thereby suppressing unfavorable coupling between spatial deformation modes. Furthermore, diaphragms can be designed in combination with longitudinal restraint systems, such as elastic displacement-restraining devices or viscous dampers, to further control the overall maximum deformation and prevent excessive displacement responses arising from insufficient stiffness in a single direction. 5.2. Main Girder Axial Force In the adopted finite element formulation, the element sectional forces are recovered from the element stiffness relationship and the corresponding displacement field. Therefore, the axial force, shear force, and bending moment reported in this study are generalized sectional internal forces obtained at the element output sections, rather than stresses at individual material points. The axial force, shear force, and bending moment responses discussed were extracted from the sectional force output of the beam elements at the representative cross sections shown in Figure 10, ensuring consistency between the finite element stress state and the reported internal-force envelopes. Axial force diagrams for typical sections of the FSWBGB, CSWBGB, and PCBGB under E1 and E2 seismic actions were calculated and are shown in Figure 14. Under both E1 and E2 seismic actions, the axial force distribution patterns of the three bridges are similar. The maximum axial force for all configurations occurs at the quarter-span section. Under E1 action, the maximum axial force of the PCBGB is approximately 5612 kN, that of the FSWBGB approximately 4510 kN, and that of the CSWBGB approximately 3906 kN. A notable increase in axial force is also observed at the quarter-span sections of the two interior spans, indicating that these regions require particular attention in seismic design. The relatively higher axial force near the central region is primarily attributed to the longitudinal restraint system of the bridge. As described in Section 3.1, the left pier and the central pier are fixed piers in the longitudinal direction, while the right pier is a movable pier. Therefore, the central pier acts not only as a vertical support but also as a major longitudinal restraint under seismic excitation. The longitudinal inertial force of the main girder is mainly transmitted to the fixed-pier system, resulting in increased axial-force demand in the girder segments adjacent to the central fixed pier and the nearby interior-span regions. In contrast, the movable pier releases part of the longitudinal constraint and consequently attracts a smaller portion of the longitudinal seismic force. Therefore, the axial-force concentration near the center of the bridge is consistent with the structural restraint mechanism rather than being an abnormal response. The CSWBGB exhibits the smallest axial force, approximately 13% lower than that of the FSWBGB. The PCBGB shows the largest axial force, approximately 43% greater than that of the CSWBGB-reflecting its relatively high self-weight and greater seismic mass participation. Under E2 action, the axial force distribution follows a similar pattern, with axial forces at typical sections approximately four times larger than those under E1. This multiplier is generally consistent with the ratio of peak ground accelerations between E2 and E1 (approximately 3.8), indicating that when the structure remains within the elastic or weakly nonlinear range, its axial force response scales approximately linearly with the seismic input level. 5.3. Main Girder Shear Force The shear force diagrams for typical sections of the CSWBGB, FSWBGB, and PCBGB under E1 and E2 seismic actions are presented in Figure 15. Under both seismic levels, the shear force distribution patterns of the three bridges are similar. Under E1 action, the maximum shear force of the PCBGB is approximately 4490 kN, that of the FSWBGB approximately 3665 kN, and that of the CSWBGB approximately 3048 kN. The maximum shear force of the CSWBGB is approximately 18% lower than that of the FSWBGB, whereas the maximum shear force of the PCBGB is approximately 22% higher than that of the CSWBGB. Under E2 action, the shear distribution follows a similar pattern, and the maximum shear forces are roughly four times those under E1 action. Notably, the shear force distribution in the CSWBGB is more uniform than that in the FSWBGB, indicating a more rational shear-sharing mechanism. This behavior originates from the high shear buckling strength of the CSWs, which enables them to stably sustain the vast majority of shear forces. Moreover, under seismic action, the CSWs dissipate a substantial amount of energy through stable plastic hysteretic deformation, effectively reducing the peak shear response and achieving a more uniform overall shear distribution. It should be noted that the maximum shear force occurs at the support sections (pier-girder joints) for all three bridge configurations. The boundary conditions at these locations are rigid monolithic connections between the pier tops and the box girder, with fixed bases at the pier bottoms. Under lateral seismic excitation, the full lateral reaction from each pier is transferred into the girder as a concentrated shear force at the rigid joint, making the support section the critical location for shear demand. 5.4. Main Girder Bending Moment The bending moment diagrams for typical sections of the three bridges under E1 and E2 seismic actions are shown in Figure 16. Under both seismic levels, the bending moment distribution patterns are similar, with peak bending moments occurring at the supports for all configurations. Under E1 action, the maximum bending moment of the PCBGB is approximately 70,010 kN·m, that of the FSWBGB approximately 42,803 kN·m, and that of the CSWBGB approximately 54,050 kN·m. The minimum bending moment occurs at the mid-span section, with a magnitude approximately 50% of the corresponding peak moment at the supports. Comparison of the bending moment envelopes reveals that, under the same seismic action, the bending moment values at a given section are approximately 20% smaller in the FSWBGB than in the CSWBGB, while those in the PCBGB are approximately 29% larger. This stable proportional relationship-the minimum mid-span bending moment being approximately 50% of the peak support moment-reflects a consistent pattern of internal force redistribution under transverse seismic loading, and may serve as a reference for simplifying the reinforcement design at mid-span sections. From the standpoint of bending moment response alone, the FSWBGB performs best among the three configurations. This is attributable to the flexural stiffness of the webs: FSWs possess greater flexural stiffness than CSWs. In the CSWBGB, the bending moment is primarily carried by the concrete top and bottom slabs, which results in lower overall flexural capacity compared to the FSW system. In this continuous rigid-frame bridge, the pier tops are rigidly connected to the box girder, and the pier bases are fully fixed. Under transverse seismic excitation, each pier acts as a vertical cantilever that transfers both its lateral reaction (as shear) and its pier-top moment (as bending moment) simultaneously into the girder at the rigid joint. Consequently, the support section must resist the combined effect of the concentrated shear force and the bending moment, making it the critical section for both internal force components. 5.5. Top Slab Normal Stress The normal stress diagrams of typical cross sections of the bridge top slab under E1 and E2 seismic actions were calculated for the CSWBGB, FSWBGB, and PCBGB, as shown in Figure 17. Under seismic action, all three bridges exhibit similar normal stress distributions, with the maximum stress occurring at the mid-span section. Under E1 action, the maximum normal tensile stress in the top slab of the PCBGB is approximately 2.50 MPa, that of the FSWBGB approximately 2.25 MPa, and that of the CSWBGB approximately 2.13 MPa. The maximum normal tensile stress in the top slab of the CSWBGB is approximately 14.8% lower than that of the PCBGB and 5.6% lower than that of the FSWBGB. Although the pattern of normal stress variation is similar among the three bridges, the overall normal stress levels in the PCBGB and FSWBGB are higher than those in the CSWBGB. This suggests that, under the same seismic action, the CSWBGB makes more efficient use of prestressing—a consequence of the low longitudinal stiffness of the CSWs, which allows prestress to be more effectively introduced into the concrete slabs. Consequently, the CSWBGB exhibits superior seismic performance compared to both the PCBGB and the FSWBGB. Under E2 action, the maximum normal tensile stress distribution follows a similar pattern, with stress magnitudes approximately three times those under E1 action. 6. Conclusions This study developed a spatial finite element model of a continuous rigid-frame bridge with corrugated steel webs (CSWBGB) using MIDAS/Civil. The seismic responses of the CSWBGB, a continuous box-girder bridge with flat steel webs (FSWBGB) of identical web thickness, and a conventional prestressed concrete box-girder bridge (PCBGB) were systematically investigated. Based on the comparative analysis of seismic performance among the three web configurations, the following conclusions are drawn: (1) Despite their distinct web types, the three bridges exhibit similar natural frequencies and mode shapes, indicating that the overall stiffness and mass distribution produce nearly equivalent dynamic effects. Notably, however, the CSWBGB possesses a relatively lower longitudinal natural frequency yet a higher transverse natural frequency compared with the FSWBGB and PCBGB. This contrasting behavior originates from the unique geometric characteristics of the CSWs-low longitudinal stiffness coupled with high out-of-plane stiffness. (2) Owing to the corrugated configuration, the CSWBGB possesses relatively low axial stiffness and consequently experiences larger longitudinal displacements than the FSWBGB. Yet, its peak axial force and shear force are significantly reduced: the axial force decreases by approximately 13% relative to the FSWBGB and by approximately 35% relative to the PCBGB, while the shear force decreases by approximately 18% and 30%, respectively. These reductions indicate that the greater axial deformability of the CSWBGB helps relieve internal force demands. Moreover, the folded geometry of the CSWs markedly enhances the transverse stiffness, achieving an optimal balance between lightweight design and structural strength, and effectively controlling lateral deformation. (3) Under E1 seismic action, the peak bending moment in the CSWBGB reaches 54,050 kN·m, approximately 20% higher than that in the FSWBGB. This difference is attributable to the inherently lower flexural stiffness of the CSWs compared with the FSWs. In the CSWBGB, the bending moment is primarily resisted by the concrete top and bottom slabs, whereas the corrugated webs contribute negligibly to the longitudinal flexural capacity-resulting in a relatively lower overall flexural resistance. (4) Under both E1 and E2 seismic actions, the peak normal tensile stress in the top slab occurs at the mid-span section for all three bridge configurations. Under E1 action, the PCBGB exhibits the highest maximum normal tensile stress at this location (approximately 2.50 MPa), whereas the CSWBGB shows the lowest (approximately 2.13 MPa), representing reductions of approximately 14.8% relative to the PCBGB and 5.6% relative to the FSWBGB. These results demonstrate that the CSWBGB experiences a more moderate stress response under seismic excitation, reflecting its superior seismic performance. This advantage stems from the accordion effect of the CSWs, which allows prestress to be more efficiently introduced into the concrete slabs, thereby improving material utilization and reducing tensile stress levels. Overall, the CSWBGB not only provides lighter self-weight than the PCBGB but also offers enhanced transverse stiffness and a more rational internal force distribution under seismic loading. Although the CSWBGB shows strong potential for practical application in high-seismicity regions, its relatively larger longitudinal displacement response should be carefully controlled in seismic design-particularly through the optimization of diaphragms and longitudinal restraint systems. Future research should systematically quantify the effectiveness of specific seismic isolation and energy dissipation devices, such as viscous dampers and isolation bearings, in controlling the longitudinal displacement of CSWBGBs, with the aim of establishing device selection and parameter optimization guidelines tailored to this bridge type. In addition, nonlinear time-history analysis incorporating material nonlinearity and geometric nonlinearity is recommended to further refine the quantitative assessment of seismic performance under rare earthquake demands. Author Contributions Conceptualization, J.H.; Methodology, B.G. and J.H.; Software, C.L.; Validation, H.P. and S.F.; Investigation, B.G.; Resources, C.L.; Data curation, C.L.; Writing—original draft, B.G.; Writing—review & editing, H.P., J.H. and S.F.; Visualization, H.P., C.L. and S.F.; Supervision, J.H.; Project administration, J.H.; Funding acquisition, H.P. All authors have read and agreed to the published version of the manuscript. Funding This research was funded by Jiangxi Provincial Natural Science Foundation, China, grant number 20252BAC240360. Data Availability Statement The original contributions presented in this study are included in the article. Further inquiries can be directed at the corresponding authors. Acknowledgments The authors gratefully acknowledge the financial support provided by the Jiangxi Provincial Natural Science Foundation, China. Conflicts of Interest Author Baojun Guo was employed by the company BCEG Civil Engineering Co., Ltd. Author Huiteng Pei was employed by the company Jiangxi Communication Design and Research Institute Co., Ltd. The remaining authors declare that the research was conducted in the absence of any commercial or financial relationships that could be construed as a potential conflict of interest. References Side elevation of box girder bridge (cm). Side elevation of box girder bridge (cm). Schematic diagram of the connections between the CSW and the top/bottom slabs. Schematic diagram of the connections between the CSW and the top/bottom slabs. Geometric parameters of box girder cross-section and CSW (m): ( a) CSWBGB; ( b) FSWBGB; ( c) PCBGB. Geometric parameters of box girder cross-section and CSW (m): ( a) CSWBGB; ( b) FSWBGB; ( c) PCBGB. Full-bridge FE model. Full-bridge FE model. Box girder cross-section: ( a) CSWBGB; ( b) FSWBGB; ( c) PCBGB. Box girder cross-section: ( a) CSWBGB; ( b) FSWBGB; ( c) PCBGB. Design acceleration response spectrum of E1 and E2 earthquake action: ( a) E1 seismic action; ( b) E2 seismic action. Design acceleration response spectrum of E1 and E2 earthquake action: ( a) E1 seismic action; ( b) E2 seismic action. PCBGB vibration modes. ( a) First-order vibration mode; ( b) Second-order vibration mode; ( c) Three-order vibration mode. PCBGB vibration modes. ( a) First-order vibration mode; ( b) Second-order vibration mode; ( c) Three-order vibration mode. CSWBGB vibration modes. ( a) First-order vibration mode; ( b) Second-order vibration mode; ( c) Three-order vibration mode. CSWBGB vibration modes. ( a) First-order vibration mode; ( b) Second-order vibration mode; ( c) Three-order vibration mode. FSWBGB vibration modes. ( a) First-order vibration mode; ( b) Second-order vibration mode; ( c) Three-order vibration mode. FSWBGB vibration modes. ( a) First-order vibration mode; ( b) Second-order vibration mode; ( c) Three-order vibration mode. Location of typical cross-sections along the main girder. Location of typical cross-sections along the main girder. Overall longitudinal deformation under E1 seismic action: ( a) CSWBGB; ( b) FSWBGB; ( c) PCBGB (mm). Overall longitudinal deformation under E1 seismic action: ( a) CSWBGB; ( b) FSWBGB; ( c) PCBGB (mm). Overall Transverse Deformation under E1 Seismic Action: ( a) CSWBGB; ( b) FSWBGB; ( c) PCBGB (mm). Overall Transverse Deformation under E1 Seismic Action: ( a) CSWBGB; ( b) FSWBGB; ( c) PCBGB (mm). Overall Vertical Deformation under E1 Seismic Action: ( a) CSWBGB; ( b) FSWBGB; ( c) PCBGB (mm). Overall Vertical Deformation under E1 Seismic Action: ( a) CSWBGB; ( b) FSWBGB; ( c) PCBGB (mm). Axial Force Distribution in Box-Girder Bridge under ( a) E1 and ( b) E2 Seismic Actions. Axial Force Distribution in Box-Girder Bridge under ( a) E1 and ( b) E2 Seismic Actions. Shear Force Distribution in Box-Girder Bridge under ( a) E1 and ( b) E2 Seismic Actions. Shear Force Distribution in Box-Girder Bridge under ( a) E1 and ( b) E2 Seismic Actions. Bending Moment Distribution in Box-Girder Bridge under ( a) E1 and ( b) E2 Seismic Actions. Bending Moment Distribution in Box-Girder Bridge under ( a) E1 and ( b) E2 Seismic Actions. Normal Stress Distribution in Box-Girder Bridge under ( a) E1 and ( b) E2 Seismic Actions. Normal Stress Distribution in Box-Girder Bridge under ( a) E1 and ( b) E2 Seismic Actions. Domestic Data on Prestressed Concrete Continuous Girder Bridges. Domestic Data on Prestressed Concrete Continuous Girder Bridges. No. Bridge Name Span Arrangement (m) Structure Form Web Thickness (cm) 1 Dongming Huanghe Bridge ୭୫ + ୭ ୍ଠ ୧୨୦ + ୭୫ Continuous rigid frame Single Room Box 40~50 2 Nanhai Haishan Bridge 660 + 120 + 66 Continuous rigid frame Single Room Box 32~60 3 Guanghe Highway Bridge 66 + 120 + 66 Continuous rigid frame Single Room Box 25~80 4 Wu Longjiang Second Bridge ୬୦ + ୩ ୍ଠ ୧୧୦ + ୬୦ Continuous beam Single Room Box 35~50 5 Xiangyin Xiangjiang Bridge ୬୫ + ୩ ୍ଠ ୧୦୦ + ୬୫ Continuous beam Single Room Box 40~50 6 Ronghua Bridge ୫୫ + ୨ ୍ଠ ୮୦ + ୪୫ Continuous beam Single Room Box 28~40 Material properties. Material properties. Component Material Elastic Modulus (MPa) Density (kN/m 3) Tensile Strength (MPa) Compressive Strength (MPa) Shear Strength (MPa) CSW Q345 206,000 78.5 310 310 180 FSW Q345 206,000 78.5 310 310 180 PC C55 35,500 25 25.3 1.96 - Peak acceleration and response spectrum of surface horizontal design motion. Peak acceleration and response spectrum of surface horizontal design motion. Transcend Probability Value Amax (gal) αmax (g) T1 (s) β mTg (s) γ 50-year 2% 220 0.651 0.1 2.9 0.55 0.9 100-year 63% 67 0.171 0.1 2.5 0.45 0.9 Cumulative modal mass participation ratios of the extracted modes. Cumulative modal mass participation ratios of the extracted modes. Bridge Type Number of Extracted Modes Longitudinal Direction (%) Transverse Direction (%) Vertical Direction (%) CSWBGB 30 91.8 90.2 92.7 FSWBGB 30 94.1 91.6 93.4 PCBGB 30 93.5 94.5 91.1 Natural vibration frequency of box girder bridge. Natural vibration frequency of box girder bridge. Modal Number Natural Vibration Frequency (HZ) CSWBGB FSWBGB PCBGB 1 0.307017 0.332972 0.323864 2 0.422982 0.391997 0.410736 3 0.613954 0.647543 0.669378 Maximum Displacement of the Main Girder under E1 and E2 Seismic Actions. Maximum Displacement of the Main Girder under E1 and E2 Seismic Actions. Bridge Type E1 Seismic Action E2 Seismic Action Long. (mm) Trans. (mm) Vert. (mm) Long. (mm) Trans. (mm) Vert. (mm) CSWBGB 10.33 6.61 1.99 47.11 30.37 8.97 FSWBGB 9.49 8.84 2.98 41.70 40.32 13.59 PCBGB 10.74 7.75 1.51 49.64 35.37 6.88 Disclaimer/Publisher’s Note: The statements, opinions and data contained in all publications are solely those of the individual author(s) and contributor(s) and not of MDPI and/or the editor(s). MDPI and/or the editor(s) disclaim responsibility for any injury to people or property resulting from any ideas, methods, instructions or products referred to in the content. © 2026 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license. Guo, B.; Pei, H.; He, J.; Luo, C.; Feng, S. Seismic Behavior of Continuous Rigid-Frame Box Girder Bridges: A Comparative Study of Different Web Configurations. Buildings 2026, 16, 2292. https://doi.org/10.3390/buildings16122292 Guo B, Pei H, He J, Luo C, Feng S. Seismic Behavior of Continuous Rigid-Frame Box Girder Bridges: A Comparative Study of Different Web Configurations. Buildings. 2026; 16(12):2292. https://doi.org/10.3390/buildings16122292 Guo, Baojun, Huiteng Pei, Jun He, Chao Luo, and Sidong Feng. 2026. "Seismic Behavior of Continuous Rigid-Frame Box Girder Bridges: A Comparative Study of Different Web Configurations" Buildings 16, no. 12: 2292. https://doi.org/10.3390/buildings16122292 Guo, B., Pei, H., He, J., Luo, C., & Feng, S. (2026). Seismic Behavior of Continuous Rigid-Frame Box Girder Bridges: A Comparative Study of Different Web Configurations. Buildings, 16(12), 2292. https://doi.org/10.3390/buildings16122292
