Seismic Response Analysis of Isolated and Nonisolated Continuous Girder Bridges under Multidimensional Near-Field Ground Motions

In this study, the responses of isolated bridges and nonisolated bridges are studied under multidimensional seismic motions. First, damage constitutive models of steel bars and concrete materials were combined with the ber beam-column element model, and the isolated bearingmodel considering bearing failure was selected.e bridge numerical analysis model was then established.e seismic responses of isolated and nonisolated bridges were analyzed under near-eld ground motions (ing-step and forwarddirectivity ground motions) and far-eld ground motions. It was found that the seismic responses of nonisolated bridges, such as deck acceleration, pier displacement, pier damage, and bearing displacement under near-eld ground motions, were higher than those under far-eld ground motions. Under far-eld ground motions, isolation bearings eectively reduced various seismic responses of structures and had isolation eects. Under forward-directivity TCU102 ground motions, the minimum isolation ratios of isolation bearings for peak acceleration of the girder in the Z direction, pier displacement, and pier shear force were −0.14, −2.65, and −0.05, respectively. e low-isolation ratio and signicant damage to the isolation bearings indicate that the isolation bearings cannot be directly used under near-eld conditions.


Introduction
ere have been numerous fault ruptures in high-density urban areas, such as the 1994 Northridge earthquake in the United States, the 1995 Kobe earthquake in Japan, the 1999 Kocaeli earthquake in Turkey, the 1999 Chi-Chi earthquake in Taiwan, and the 2008 Wenchuan earthquake in Sichuan. ese earthquakes caused extensive building and bridge damage throughout the cities. e concentration of earthquake damage is re ected in the extreme earthquake area, enlisting the need for further study and development of better engineering practice. Near-eld motion characteristics [1] mainly include forward-directivity, ing-step, hanging-wall, and signi cant vertical ground motion. Among them, the forward-directivity e ect is one of the main reasons causing the characteristics of near-eld ground motions' pulse, which is caused by the Doppler e ect of fault rupture propagation and shows a bidirectional velocity pulse [2]. e ing-step e ect is when the two plates of the fault are relatively dislocated or slide during an earthquake, andnally, permanent ground displacement is generated in the sliding direction [3]. e velocity pulse caused by the ingstep e ect is related to the magnitude of permanent displacement and the timing of permanent displacement, which is a unidirectional pulse. Peak pulse velocity has a signi cant in uence on structural demands [4,5]. Velocity pulse is one of the leading causes of near-eld structure failure. Under near-eld pulse-like ground motions, the e ectiveness and applicability of isolation technology need to be studied and globally implemented [1].
Bridge isolation technology has been widely used in the seismic design of bridges. However, when designing and analyzing isolated bridges, the in uence of near-eld ground motions on the seismic responses of bridges is often ignored.
With the in-depth study of near-field ground motions, the characteristics, laws, and potential failure mechanism of seismic response of isolated bridges under near-field ground motions have gradually become the focus of research. Jonsson et al. [6] analyzed the response of an isolated bridge under strong near-field ground motions and found that a practical design of an isolation system can effectively prevent the bridge structure from damage. Wei et al. [7] analyzed the seismic response of an isolated, simply supported girder bridge and found that the isolated bearing still has some seismic absorption effect under near-field ground motions. Jalali et al. [8,9] analyzed the response of a three-span bridge with mid-span isolation under near-field pulse-like ground motion. ey studied the difference between parallel fault and vertical fault ground motion to the structural response. Ismail et al. [10] analyzed and studied the isolation effect of a new type of roll-in-cage (RNC) isolator under near-field, long-period, and pulse-like ground motions and verified the effectiveness of the isolation bearing. Losanno et al. [11] optimized the damping parameters of isolated bridges through numerical simulation and compared the influence of near-field and far-field ground motions on isolation effects. Liao et al. [12] found that the PGV/PGA ratio of nearfield ground motion greatly influences the bridge response. Kalkan and Kunnath [13] studied the effect of forward directivity and fling step on the seismic response of steel frame structures. Zheng et al. [14] analyzed the seismic performance of bridges installed with a sliding-lead rubber-bearing isolation system subjected to near-fault earthquakes. Jiang et al. [15] proposed a risk-based approach to study the pulse effect on the isolator optimization of bridges in near-fault zones. In the study of isolated bridges, the influence of forward directivity and fling step on bridge response is seldom distinguished. In addition, most studies only consider unidirectional or bidirectional ground motion, but not three-dimensional ground motion.
At present, the failure of isolated bearings is not considered in the seismic response research of most isolated bridges, or only the critical displacement is used to judge if the bearings are damaged, which is unsafe. Buckle et al. [16] found that the critical load value of the bearing decreases with the increase in lateral displacement through experimental research. It is essential to consider the change in critical load when designing bearings. Li et al. [17] proposed a three-dimensional isolation bearing simulation model that considers the horizontal bidirectional coupling effect and vertical stability, which can better simulate the nonlinear mechanical characteristics and vertical stability of isolation bearings. e research on isolated bridges typically does not focus on pier damage because the isolated bearings can better protect the safety of piers under far-field ground motions, and the piers are basically in an elastic state with minor damage. However, under near-field ground motions, piers may be damaged. Zhong et al. [18] proposed an uncoupled multivariate power model to estimate the performancebased seismic damage states of column curvature ductility.
e fiber-element model with computational efficiency and numerical accuracy is suitable for the nonlinear analysis of reinforced concrete members [19]. Heo and Kunnath [20] proposed a damage model of reinforced concrete members based on material damage using the fiber-element model. e model takes the damage indexes of the critical fibers of compressed concrete and the critical fibers of reinforcement in the core area of concrete as the cross-sectional damage indexes. Li et al. [21] and Gao et al. [22] used the fiberelement model combined with the Faria-Oliver uniaxial concrete damage model [23] to analyze the damage to reinforced concrete members.
In this study, based on the refined simulation and analysis platform for structures (RSAPS) [24] previously developed, the fiber beam-column element model and the three-dimensional isolation bearing model [17] are used to simulate the nonlinear characteristics of piers and isolation bearings, respectively.
e concrete damage constitutive model [25] and the steel bar constitutive model [25] were combined with the fiber beam-column element, establishing the damage analysis model of the reinforced concrete bridge.
ree-dimensional near-field ground motions (fling-step and forward-directivity ground motions) and far-field ground motions were selected to analyze the seismic response of isolated bridges and nonisolated bridges, respectively.

Ground Motion Record Selection
According to the research results of Kalkan and Kunnath [13], Sehhati et al. [26], Moniri [27], Li et al. [28], and Wang and Bai [29], seven forward-directivity ground motions, seven fling-step ground motions, and seven far-field ground motions were obtained from the Pacific Earthquake Engineering Research Center database [30]. e selected ground motion records and their parameters are shown in Tables 1-3.

Finite Element Model of Reinforced Concrete Bridge
In this study, the simulation analysis platform RSAPS [24] was used to simulate and analyze the seismic response of reinforced concrete bridges. e RSAPS platform was established based on the subroutine (UEL) interface of the general finite element software ABAQUS and mainly included the fiber beam-column element model, isolation element model, and various material constitutive models of concrete and steel. Good simulation results [24] have been achieved for static and dynamic nonlinear behavior simulation of reinforced concrete members. e specific bridge model is given below.

Analysis Model of the Bridge.
e seismic response analysis of a bridge ( Figure 1) in reference [24] was carried out. e bridge is a continuous girder bridge with a span of 30 m and five spans. e girder has a single-box and threechamber section with a height of 1.88 m and a width of 8.025 m. e pier height is 6.6 m, and the diameter is 1.219 m. e section of the girder and the pier is shown in Figure 2.
In modeling analysis, because the main bridge girders are mostly prestressed reinforced concrete structures, the girders seldom undergo plastic deformation or damage behavior when subjected to an earthquake. e linear elastic beam element was used to simulate the girders, and each span was divided into 15 elastic beam elements. e pier was the primary stress member; strong nonlinear behaviors such as plasticity and damage appear during an earthquake; and they were simulated using the fiber beam-column elements.
For the isolated bridge in reference [24], LRB700-140 isolation bearings were selected. e bearing diameter is 700 mm, the lead diameter is 140 mm, the total thickness of the rubber layer is 110 mm, and the total thickness of the steel plate layer is 75 mm. e preyield stiffness of the bearing is 17.771 kN/mm, the postyield stiffness is 1.367 kN/ mm, and the yield force is 94.2 kN. e critical design load is 4618 kN. e LRB bearings were simulated by isolation elements (shown in Section 3.3).
To study the isolation effect of LRB isolation bearings, the bearings in the nonisolated bridge model were assumed to be ordinary linear elastic bearings, and the stiffness of the bearings was taken as the preyield stiffness of LRB bearings at corresponding positions in the isolated bridges [24]. e bearings in nonisolated bridges were simulated as spring elements.

Pier Analysis Model.
e fiber beam-column element was adopted for modeling the nonlinear behavior of a pier. Each member was divided into six fiber beam-column elements along with the height. Each element adopts four Gauss-Lobatto integral sections. Each section was divided into 48 longitudinal reinforcements and 216 concrete fibers, including 180 core and 36 protective layer concrete fibers. e cross-sectional fiber discretization method is shown in Figure 3.

Constitutive Model for Concrete Material.
Yassin's uniaxial concrete model [31] was implemented in this study (as shown in Figure 4). Yassin's model has the following advantages: (1) this model can simulate the continuous stiffness degradation effect during unloading and reloading with an increase in concrete compressive strain; (2) under repeated loading and unloading, the hysteretic performance of concrete materials can be effectively simulated. e model can also simulate the stirrup constraint effect by modifying the characteristic parameters of the concrete materials.   To consider the damage performance in the uniaxial constitutive model of concrete, the tensile and compressive damage indexes of concrete are used to describe the tensile and compressive damage of concrete, respectively. e damage distribution and its evolution process can be intuitively described.
(1) Compression Damage. According to the basic principle of damage mechanics and the characteristics of the concrete constitutive model, the calculation method of compression damage index is as follows [25]: where D c is the compression damage index of concrete; is the initial compression damage module, determined according to the compression damage starting point D 0 (ε c c d0 , σ c c d0 ) and the focus R(ε r , σ r ); and E cm � (σ c cm − σ r )/(ε c cm − ε r ) is the module of the current unloading point, determined according to the unloading point D(ε c cm , σ c cm ) and the focus R(ε r , σ r ). Compression damage of concrete only occurs during loading but does not occur during unloading. New damage occurs only after reloading reaches the previous unloading point. According to Yue et al.'s research [32], a point when stress in the rising section of the skeleton curve reaches 0.3 f c ′ is selected as the damage starting point. e stress-strain relationship and corresponding compression damage index of concrete under compression are shown in Figure 5.
For the convenience of expression and understanding, the stress is placed in the first quadrant in the introduction of the compression characteristics of the above concrete, and the principle of positive tension and negative compression is followed in practical application.
(2) Tensile Properties. A linear model was adopted for the stress-strain relationship curve of the tensile skeleton, as shown in Figure 6. f t is the peak tensile stress; ε cr is the strain corresponding to the peak tensile stress; E c0 is the initial tangent modulus; and ε ut is the ultimate tensile strain. e tensile damage index is calculated as follows [25]: where E tm is the secant modulus of the current unloading point; ε t cm is the strain at the current unloading point. Tensile concrete damage only occurs during loading and does not occur during unloading. New damage occurs only after reloading reaches the previous unloading point. e stress-strain relationship and corresponding tensile damage index of concrete under tension are shown in Figure 7. e parameters of the concrete material are shown in Table 4.

Constitutive Model for Steel Material.
In the analysis of this study, the stress-strain relationship of steel bars was expressed by the modified Menegotto-Pinto constitutive model [33]. e model was proposed by Menegotto and Pinto and modified by Filippou et al. [34] to consider the influence of the isotropic strengthening effect.
e Bauschinger effect under cyclic loads is considered in the model, is in good agreement with experimental results, and has high solution e ciency. Overall, this model has been extensively used. e Bonora damage model [35] was introduced into the modi ed Menegotto-Pinto constitutive model to consider the damage and fracture behavior of steel bars. e Bonora damage model is an elastic-plastic damage constitutive model based on continuous damage mechanics. e model adopts nonlinear damage evolution criteria, which can better simulate the damage performance of steel. Compared with the test, the model has high simulation accuracy. e damage index in the Bonora damage model is calculated as follows [35]: where _ D is the damage increment; D is the cumulative damage value; D 0 is the initial damage value; D cr is the critical damage value; ε cr is the critical strain corresponding to the critical damage value; ε th is the threshold strain for the start of damage; dp is equivalent plastic strain increment; p is equivalent plastic strain; α is the damage parameter; and f(σ m /σ eq ) is the in uence factor in the triaxial stress state and is taken as 1 for the uniaxial constitutive model. e parameters used in this study are selected according to reference [35].
To account for the occurrence of repeated tensioncompression overloads causing failure, Pirondi and Bonora [36] modi ed the Bonora damage model. It is considered that the steel bar is damaged only when it is tensile, so only the e ect of tensile plastic strain is considered when calculating the damage index. Figures 8 and 9 are the stress-   e parameters of the steel material are shown in Table 5.

Force-Displacement Response of a Pier.
e forcedisplacement response of a pier under a constant axial force of 2889.24 kN is shown in Figure 10. e horizontal bearing capacity of the pier is 1241.8 kN.

Bearing Analysis Model.
Li et al. [17] developed an LRB isolation element model based on ABAQUS, which adopted the bidirectional coupled Bouc-Wen model improved by Casciati [37] in the horizontal direction. e restoring force was combined with the following relationship: where F 1 and F 2 are the restoring forces of the lead rubber bearings in the X and the Y directions, respectively; U 1 and U 2 represent the relative displacement of the lead rubber bearing in the X and the Y directions, respectively; α is the ratio of postyield-to-preyield sti ness; k b is the initial sti ness; c b is viscous damping of the lead rubber bearings; and Z 1 and Z 2 are the hysteretic displacements in the X and the Y directions, respectively, satisfying the following relationship: where U y is the yield displacement of the lead rubber bearing; A, c, and β are the parameters that control the shape and size of the restoring force-displacement hysteresis loop of the lead rubber bearing, generally taking 1, 0.5, and 0.5, respectively; and sgn is a symbolic function. e overlap area method [38] was used to determine the bearing capacity (critical load) of the bearing under a given lateral displacement as follows: where P cr ′ is the bearing capacity (critical load) of a lead rubber bearing considering the in uence of lateral displacement; A r is the area of the overlapping part of the upper and lower sections of the bearing; A b is the cross-sectional area of the lead rubber bearing; and P cr is the bearing Mathematical Problems in Engineering capacity (critical load) of the lead rubber bearing without lateral displacement. e LRB isolation element model is shown in Figure 11.

Seismic Response Analysis of Bridges
Seismic responses of isolated and nonisolated bridges were analyzed, respectively, and the seismic responses of bridge structures under three di erent ground motions were studied. e spatial di erential e ect of ground motion was not considered in the analysis, and uniform excitation was used for ground motion input. e PGA in the X direction (X direction (longitudinal direction) of the bridge) of the selected ground motion record was uniformly adjusted to 0.4 g to correspond to a rare earthquake of 8 degrees [39]. e PGA in the Y direction (Z direction (transversal direction) of the bridge) was uniformly adjusted to 0.34 g. e PGA in the Z direction (Y direction (vertical direction) of the bridge) was uniformly adjusted to 0.26 g. Figures 12-14, respectively, show the peak acceleration responses of the girders at the top of the bearings of nonisolated bridges and isolated bridges under di erent types of ground motions. Overall, the peak accelerations at all positions of nonisolated bridges and isolated bridges were the same under the same ground motion in the X direction, regardless of the ing-step ground motions, forward-directivity ground motions, or far-eld ground motions. Under the same ground motion in the Y direction, the peak acceleration at each position of the nonisolated bridge had a W-shaped distribution, while the peak acceleration at each position of the isolated bridge was the same. Under the di erent ground motions, the peak acceleration response of the bridges in the Z direction was greater than that in the X direction in most conditions, although the peak acceleration in the X direction was more signi cant when ground motions were input. is was more obvious at the end of nonisolated bridges. e X-direction acceleration of the B02 position of the nonisolated bridge     For the ing-step ground motions, under TCU068 ground motion, the isolation e ect of the isolation bearings was the least in the X direction and Z direction, especially in the X direction, with the lowest isolation ratio being only 0.06. is indicates that the isolation bearings hardly a ected the structure concerning seismic isolation. Except under the    relatively poor, with isolation ratios of ∼0.4. Under other ground motions, the isolation e ects of the isolation bearings increased, with the isolation ratios all greater than 0.6. e isolation ratios of the isolation bearings were greater than 0.6 under only three ground motions in the Z direction.

Girder Acceleration.
Under the TCU102 ground motion, the isolation e ect of the isolation bearings was the least. Even at the B4 position, the isolation ratio was −0.14, indicating that the isolation bearing did not reduce the acceleration response; instead, it increased the acceleration response.
is shows that the  isolation bearings cannot be used under TCU102 ground motion. Under the far-field ground motions, the isolation effects of the isolation bearings were excellent in the X direction. e isolation ratios were all above 0.8; the isolation bearings had a good isolation effect in the Z direction, and the lowest isolation ratio was still greater than 0.5. Figures 15-17 show the peak displacement of piers of nonisolated bridges and isolated bridges under different types of ground motions, respectively. e displacement of the pier was calculated by ������� � D 2 X + D 2 Z (where D X and D Z are the displacements of the pier in the X and Z directions, respectively). e peak displacements of the middle piers of the nonisolated bridges were slightly greater than that of the side piers under the same ground motion, regardless of the fling-step ground motions, forward-directivity ground motions, or far-field motions. However, the peak displacements of the isolated bridge piers were the same. Under the fling-step ground motions, the maximum peak displacement and the minimum peak displacement of nonisolated bridge piers were 219.49 mm and 95.86 mm, respectively. e maximum peak displacement of isolated bridge piers was 98.40 mm, and the minimum peak displacement was 7.42 mm. Under the forward-directivity ground motions, the maximum peak displacement of nonisolated bridge piers was 271.67 mm, and the minimum peak displacement was 77.95 mm. e maximum peak displacement of the isolated bridge piers was 715.49 mm, and the minimum peak displacement was 9.91 mm. Under the far-field ground motions, the maximum peak displacement of nonisolated bridge piers was 169.75 mm, and the minimum peak displacement was 39.07 mm. e maximum peak displacement of the isolated bridge piers was 8.75 mm, and the minimum peak displacement was 2.65 mm. e maximum peak displacement and minimum peak displacement of the nonisolated bridge piers under fling-step ground motion and forward-directivity ground motion were more extensive than those under far-field ground motion. e maximum peak displacement of the isolated bridge piers under fling-step ground motion and the forward-directivity ground motion was much larger than that under far-field ground motion, especially under forward-directivity ground motion. e minimum peak displacement of the isolated bridge piers under fling-step ground motion and the forward-directivity ground motion was much larger than that under far-field ground motions, but its values were overall small.

Pier Displacement.
For the fling-step ground motions, under TCU068 ground motion, the isolation effect of the isolation bearings was very poor, with the lowest isolation ratio of only 0. 13. is indicates that the isolation bearings have little influence on seismic isolation. Under other ground motions, the isolation effects of the isolation bearings were better, with ratios all above 0.7 and a maximum of 0.96. For the forwarddirectivity ground motions, except under TCU102 ground motion, the isolation effects of isolation bearings under other ground motions were better, with the isolation ratios all greater than 0.7. e isolation bearings do not affect seismic isolation under TCU102 ground motion. e isolation ratio of each position was negative, below −1.86, with a minimum of −2.65. is shows that the isolation bearings do not reduce the displacement response of piers but significantly increase the displacement response of piers. e isolation bearings cannot be used under TCU102 ground motion. Under the far-field ground motions, the isolation effects of isolation bearings were outstanding, and the isolation ratios were all above 0.9. Figures 18-20 show the peak shear force of piers of nonisolated bridges and isolated bridges under different ground motions. e peak shear force of each pier of the nonisolated bridge reached the maximum bearing capacity of the pier of ∼ 1200 kN under the fling-step ground motions, forward-directivity ground motions, or far-field ground motions. e difference in the maximum bearing capacity of piers is caused by a difference in the vertical loads on the piers. For the fling-step ground motions, under the TCU072 ground motion, the isolation bearings significantly reduce the peak shear force of the piers, with a minimum value of 352.25 kN. Under the TCU068 ground motion, the isolation bearings did not reduce the peak shear force of the piers, with the peak shear force of the piers above 1200 kN. For the forward-directivity ground motions, except under JFPABuilding and TCU102 ground motions, the isolation bearings significantly reduced the peak shear force of the piers under other ground motions, and their values were between 400 kN and 600 kN. Under JFPABuilding and TCU102 ground motions, the isolation bearings did not reduce the peak shear force of each pier, with the peak shear force of each pier ∼1200 kN. Under the far-field ground motions, the isolation bearings significantly reduced the peak shear force of piers, and the peak shear force of each pier was below 400 kN. e isolation effect of the isolation bearing under TCU068 fling-step ground motion was minor, with the lowest isolation ratio being −0.01.

Pier Shear Force.
is indicates that the isolation bearings do not play roles in seismic isolation. e isolation bearings under the TCU068 ground motion were not applicable. e isolation effect of the isolation bearing under the Yarimca ground motion was also poor, with the isolation ratios all less than 0.2. e isolation effects of the isolation bearings under other ground motions increased with ratios ranging between 0.4 and 0.75. For the forwarddirectivity ground motion, the isolation effect of the isolation bearing under TCU102 ground motion was poor, with the lowest isolation ratio being −0.05. is indicates that the isolation bearings do not contribute to seismic isolation. e isolation bearing under TCU102 ground motion was not applicable. e isolation effect of the bearings under the JFPABuilding ground motion was poor, with the isolation ratios all less than 0.1. e isolation effects of the isolation bearings under other ground motions were improved, with isolation ratios between 0.5 and 0.7. Under the far-field ground motions, the isolation effects of the isolation bearings were excellent, with isolation ratios all above 0.7.

Pier Damage.
Under various conditions, the reinforcement damage of each pier was either negligible or nonexistent. e tensile damage index of concrete quickly reached 1; therefore, only the compression damage of concrete in the core area was analyzed. Based on the method Heo and Kunnath [20] proposed for estimating the damage index of the concrete member, the damage index of the pier was estimated using the damage index of the most critical bers for concrete in the core area and reinforcing steel. Because the reinforcement damage of each pier was very   small or none, it was reasonable to consider the damage index of the most critical ber for concrete in the core area as the representative damage index of the pier. Figures 21-23, respectively, show the peak damage index of the compression concrete in the pier core area of nonisolated bridges and isolated bridges under di erent ground motions. Under the same ground motion, the peak damage indexes of the middle piers were slightly greater than that of the side piers for the nonisolated bridge, regardless of the ing-step ground motions, forward-directivity ground    motions, or far-eld motions. However, the peak damage index to each pier of the isolated bridge was the same. is is consistent with the peak displacement response of piers.
Under the TCU072 and TCU076 ing-step ground motions, the peak damage index of each pier of the nonisolated bridge reached 1; the concrete in the core area of each pier was crushed. Under TCU054, TCU065, and TCU068 ground motions, the peak damage index of the middle piers (P2 and P3) of the nonisolated bridge also reached 1, indicating that the area was also destroyed. Under Abbar-non isolated Abbar-isolated AnzaTule-non isolated AnzaTule-isolated BakerFire-non isolated BakerFire-isolated BigBear-non isolated BigBear-isolated Calexico-non isolated Calexico-isolated LACCNorth-non isolated LACCNorth-isolated Ta -non isolated Ta -isolated  the Yarimca ground motion, the peak damage index of each pier of the nonisolated bridge did not reach 1 but was greater than 0.85, indicating extensive damage to the pier. Under the TCU082 ground motion, the peak damage index of each pier of the nonisolated bridge was relatively small, ranging from 0.5 to 0.8. Under the TCU068 ground motion, the peak damage indexes of the isolated bridge ranged between 0.5 and 0. 8  For the forward-directivity ground motions, although the peak damage indexes of the piers (P2 and P3) of the nonisolated bridge under Duzce and TCU101 ground motions did not reach 1, they were greater than 0.85, showing signi cant damage to the piers. e peak damage indexes of the side piers (P1 and P4) were relatively small, but they still varied between 0.45 and 0.8. Under other ground motions, the peak damage index of each pier of the nonisolated bridge reached 1, indicating complete destruction to the core. Under the TCU102 ground motion, the peak damage index of each isolated bridge pier reached 1, indicating that the concrete in the core area of each pier was crushed, and the isolation bearings do not play a role in reducing the damage to the pier. Under the JFPABuilding ground motion, the peak damage index of each isolated bridge pier varied between 0.3 and 0.35, and the damage to each pier was relatively minor. Under other ground motions, the peak damage index of each isolated bridge pier was less than 0.03, and the damage to each pier was minimal.
Under the BigBear and Taft far-eld ground motions, the peak damage indexes of the middle piers (P2 and P3) of the nonisolated bridge reached 1, and the concrete core area was crushed. Although the peak damage indexes of the side piers (P1 and P4) did not reach 1, there were still above 0.8, indicating extensive damage to the piers. Under the BakerFire ground motion, the peak damage index of pier P3 of the nonisolated bridge reached 1, and the concrete in the core area of the pier was crushed. e peak damage index of pier P2 reached 0.97. Although the concrete in the core area of pier P2 was not crushed, it was extensively damaged. e peak damage indexes of the side piers (P1 and P4) ranged between 0.3 and 0.45, and the damage was relatively small. Under the LaCCNth ground motion, the peak damage indexes of the middle piers (P2 and P3) of the nonisolated bridge were greater than 0.9, and the peak damage indexes of the side piers (P1 and P4) were above 0.7. Each pier has extensive damage. Under other ground motions, the peak damage index of each pier of the nonisolated bridge was less than 0.7, with a minimum value of 0.18; here, the pier damage is relatively minor. Under the far-eld ground motions, the peak damage index of each pier of the isolated bridge was 0, and there was no compression damage to the concrete in the core area of each pier. e lowest isolation ratio under the TCU068 ing-step ground motion was only 0.06. Here, the isolation e ects of the isolation bearings were minimal, and the isolation bearings hardly a ected the seismic isolation of the bridge. Under other ground motions, the isolation ratios were greater than 0.8, and the maximum even reached 1. e isolation bearings were e ective under these ground motions. For the forward-directivity ground motions, the isolation ratio of each isolation bearing under the TCU102 ground motion was 0, and the isolation bearings did not in uence seismic isolation. e isolation ratios under the JFPABuilding ground motion were between 0.65 and 0.7, and the isolation e ects of the isolation bearings were moderate. e isolation ratios under other ground motions were greater than 0.95, and the isolation e ects of isolation bearings were signi cant. Under far-eld ground motions, all the isolation ratios reached 1, and isolation bearings were very e ective.   and 172.08 mm, respectively. e peak displacements of the end bearings of the nonisolated bridges were much larger than those of the middle bearings. e peak displacements of all bearings of the isolated bridges were the same under the same ground motion.
For the ing-step ground motions, the peak displacements of the end bearings (Bearing01 and Bearing02) of the isolated bridge under the TCU072 ground motion were slightly smaller than those of the nonisolated bridge. e peak displacements of the middle bearings (Bear-ing1∼Bearing4) of the isolated bridge were much larger than    TCU054  TCU065  TCU068  TCU072  TCU076  TCU082  Yarimca  Bearing01  1  0  1  0  0  1  1  Bearing1  1  1  1  0  1  1  1  Bearing2  1  1  1  1  1  1  1  Bearing3  1  1  1  1  1  1  1  Bearing4  1  1  1  1  1  1  1  Bearing02  0  0  1  0  0  1  1  Forward-directional ground motions  Bearing  Duzce  JFPABuilding  Lexington Dam  SCSE  SOVMFF  TCU101  TCU102  Bearing01  0  1  0  0  0  0  1  Bearing1  1  1  1  1  1  1  1  Bearing2  1  1  1  1  1  1  1  Bearing3  1  1  1  1  1  1  1  Bearing4  1  1  1  1  1  1  1  Bearing02  0  1  Under other ground motions, the peak displacement of the isolated bridge bearings ranged between 200 mm and 400 mm. For the far-field ground motions, the peak displacement of each isolated bridge bearing was smaller than that of the nonisolated bridge under the AnzaTule ground motion. Under the BigBear, Calexico, and Taft ground motions, the peak displacement of the end bearings of the isolated bridges (Bearing01 and Bearing02) was slightly less than that of nonisolated bridges. Under other ground motions, the peak displacement of the bearings of the isolated bridges was more significant than that of nonisolated bridges. However, the maximum peak displacement of the bearings of the isolated bridges was only 174.23 mm. is value was slightly larger than the minimum value of the bearing peak displacement of the isolated bridge under the fling-step ground motion, reaching 169.98 mm, and smaller than the minimum value of 227.18 mm for the bearing peak displacement of the isolated bridge under the forward-directivity ground motion.

Failure of Isolation
Bearings. e failure of isolation bearings under different types of ground motions is given in Table 6. Here, 0 indicates that the isolation bearing was not damaged, and 1 indicates that the isolation bearing was damaged. e failure index of the bearing model was only used for identification in the analysis, and the mechanical properties of the bearing were not affected. Most of the isolation bearings were damaged under the ground motion of the fling step and forward directivity; only 8 and 10 isolation bearings were not damaged, respectively. All isolation bearings were not damaged under the far-field ground motions. Overall, this shows that an isolation bearing cannot be directly used under near-field conditions.

Discussion
is study found that the isolation bearings can effectively reduce various seismic responses of the bridge under the farfield ground motions. Under the near-field ground motions, the isolation bearings still have a specific isolation effect in most cases, consistent with Jonsson et al.'s [6] and Wei et al.'s [7] research results. At the same time, it was found that the isolation bearings do not work and even amplified the seismic response of the bridge, such as under the flingstep TCU068 ground motion and the TCU102 forwarddirectivity ground motion. In addition, extensive damage to the isolation bearing under near-field ground motions indicates that the isolation bearings cannot be used in nearfield conditions.
In the analysis, the differential effect of the ground motion is not considered, which may have a particular impact on the seismic response of a bridge. In addition, the study of near-field ground motion characteristics also needs to be further strengthened to better study its influence on the seismic response of a bridge.

Conclusions
is study created the damage constitutive models of steel and concrete material, the fiber beam-column element model, and the bearing model to establish a numerical analysis model of a bridge to study the effects of near-field (fling-step and forward-directivity ground motions) and farfield ground motions on isolated and nonisolated bridges. e following research results are obtained: (1) e seismic responses of nonisolated bridges, such as girder acceleration, pier displacement, pier damage, and bearing displacement under fling-step ground motions and forward-directivity ground motions, increased to a certain extent compared to those under far-field ground motions. e pier displacement of the nonisolated bridges under forward-directivity ground motions is greater than that under fling-step ground motions, which also leads to more significant pier damage of the nonisolated bridges under forward-directivity ground motions.
(2) Isolation bearings can effectively reduce various seismic responses of a bridge (girder acceleration, pier displacement, pier shear force, and pier damage) under far-field ground motions. Isolation bearings can effectively isolate and protect the bridge. Under the TCU068 fling-step ground motions and TCU102 forward-directivity ground motions, the isolation bearings do not work and even amplify the seismic response of the bridge. However, the isolation bearings have an excellent isolation effect on the seismic response of bridges under most near-field ground motions. e isolation effect of isolation bearings under forward-directivity ground motion is better than that under fling-step ground motion.
(3) Compared with the far-field ground motions, the displacement of the isolation bearings under the fling-step ground motion and the forward-directivity ground motion is more extensive. is is the main reason for the damage to isolation bearings. Extensive damage to the isolation bearings under the near-field ground motions indicates that they cannot be directly used under near-field conditions.
(4) Under the near-field ground motions, nonisolated bridges often fail due to severe damage to the piers, while isolated bridges often fail due to the destruction of isolation bearings; sometimes, even the setting of the isolation bearings aggravates the seismic response of the bridges and leads to bridge failure. Under the far-field ground motions, the nonisolated bridge may not be seriously damaged, or the bridge may fail due to the severe damage to the piers. e setting of the isolation bearings can effectively reduce the response of the bridge and ensure the safety of the bridge.
ough the research results in this study cannot be directly applied in design practice, it was found that the isolation bearing may be damaged under near-field ground motions and cannot be directly implemented in practice, which needs further study. In addition, it was found that the failure modes of the isolated and nonisolated bridges under near-field ground motions were different, which provides an essential basis for the seismic design of bridges.

Data Availability
e data used to support the findings of this study are available from the corresponding author upon request.

Conflicts of Interest
e authors declare that there are no conflicts of interest regarding the publication of this paper.