SEISMIC PERFORMANCE OF SMA RETROFITTED MULTIPLE-FRAME RC BRIDGES SUBJECTED TO STRONG MAIN SHOCK-AFTERSHOCK SEQUENCES

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1 10NCEE Tenth U.S. National Conference on Earthquake Engineering Frontiers of Earthquake Engineering July 21-25, 2014 Anchorage, Alaska SEISMIC PERFORMANCE OF SMA RETROFITTED MULTIPLE-FRAME RC BRIDGES SUBJECTED TO STRONG MAIN SHOCK-AFTERSHOCK SEQUENCES W. Huang 1 and B. Andrawes 2 ABSTRACT Using Shape Memory Alloy (SMA) wires to retrofit vulnerable reinforced concrete (RC) bridge piers had shown promising results in recent analytical and experimental researches. However, most of previous researches had focused on performance of retrofitted RC piers on element level while the global response of SMA-retrofitted RC bridge has not yet been studied. Therefore, the main objective of this numerical study is to investigate the global seismic response of full-scaled SMA-retrofitted multiple-frame RC bridges, especially focusing on the sequence of formation of plastic hinges at bridge piers. Furthermore, the bridge is subjected to a sequence of strong main shock and aftershock ground motions to determine the effect of sequential seismic loadings on retrofitted bridge piers and on other bridge components such as expansion joints and abutments. The results demonstrate the level of retrofitting needed to improve the resilience of the bridge to a given target performance, which is measured by the aftershock PGA the bridge can resist. In addition, the optimal level of retrofit for the bridge is identified. It is also found that there is an increase in demands on intermediate hinges and abutments as retrofitted bridge is subjected to increasing intensity of aftershock. 1 Graduate Research Assistant, address: 3109 Newmark Civil Engineering Lab, 205 North Matthews Ave., Urbana, IL 61801, USA. whuang41@illinois.edu 2 Associate Professor, address: 3122 Newmark Civil Engineering Lab, 205 North Matthews Ave., Urbana, IL 61801, USA. Phone: (217) andrawes@illinois.edu Huang W, Andrawes B. Seismic performance of SMA retrofitted multiple-frame RC bridges subjected to strong main shock-aftershock sequences. Proceedings of the 10 th National Conference in Earthquake Engineering, Earthquake Engineering Research Institute, Anchorage, AK, 2014.

2 10NCEE Tenth U.S. National Conference on Earthquake Engineering Frontiers of Earthquake Engineering July 21-25, 2014 Anchorage, Alaska Seismic performance of SMA retrofitted multiple-frame RC bridges subjected to strong main shock-aftershock sequences W. Huang 1 and B. Andrawes 2 ABSTRACT Using Shape Memory Alloy (SMA) wires to retrofit vulnerable reinforced concrete (RC) bridge piers had shown promising results in recent analytical and experimental researches. However, most of previous researches had focused on performance of retrofitted RC piers on element level while the global response of SMA-retrofitted RC bridge has not yet been studied. Therefore, the main objective of this numerical study is to investigate the global seismic response of full-scaled SMA-retrofitted multiple-frame RC bridges, especially focusing on the sequence of formation of plastic hinges at bridge piers. Furthermore, the bridge is subjected to a sequence of strong main shock and aftershock ground motions to determine the effect of sequential seismic loadings on retrofitted bridge piers and on other bridge components such as expansion joints and abutments. The results demonstrate the level of retrofitting needed to improve the resilience of the bridge to a given target performance, which is measured by the aftershock PGA the bridge can resist. In addition, the optimal level of retrofit for the bridge is identified. It is also found that there is an increase in demands on intermediate hinges and abutments as retrofitted bridge is subjected to increasing intensity of aftershock. Introduction Recent analytical and experimental studies have shown that applying active confinement through forms of Shape Memory Alloy (SMA) wire is an effective retrofit method for vulnerable reinforced concrete (RC) columns. Prestrained SMA spirals are wrapped and anchored around the RC columns and heated to exert the active confining pressure in the form of hoop stress. Shin and Andrawes [1] had shown that the active confinement and passive confinement (through dilation of concrete column) could increase flexural ductility of RC bridge column substantially, and consequently, improve the seismic performance of the columns. However, previous researches had mainly focused on assessing the performance of retrofitted bridge columns on the component level. The global behavior of bridge after retrofitting its one or more components has not yet been studied. Moreover, the change of demands on other components of bridges (i.e. bearings, abutment, or intermediate hinges) after retrofitting needs to be investigated to avoid local failures and to ensure the overall integrity of the bridge. 1 Graduate Research Assistant, address: 3109 Newmark Civil Engineering Lab, 205 North Matthews Ave., Urbana, IL 61801, USA. whuang41@illinois.edu 2 Associate Professor, address: 3122 Newmark Civil Engineering Lab, 205 North Matthews Ave., Urbana, IL 61801, USA. Phone: (217) andrawes@illinois.edu Huang W, Andrawes B. Seismic performance of SMA retrofitted multiple-frame RC bridges subjected to strong main shock-aftershock sequences. Proceedings of the 10 th National Conference in Earthquake Engineering, Earthquake Engineering Research Institute, Anchorage, AK, 2014.

3 This paper aims at studying the efficacy of the new SMA retrofit technique in mitigating the damage sustained by bridges under strong earthquakes, which are often characterized by sequence of earthquakes in the form of strong main shock followed by multiple strong-tomoderate aftershocks. Recent strong earthquakes, namely Chi-Chi (1999), Tohoku (2011), and Christchurch (2011), are good examples of strong earthquake followed by a large numbers of aftershocks. Nevertheless, past and current structural design codes overlook the potential danger of aftershocks on already damaged structures by only considering earthquake as a single event of loading on structures. Early analytical investigation on effects of aftershocks on structural response focused on single-degree-of-freedom (SDOF) system with elastic or elastic-perfectlyplastic (EPP) inelastic systems using artificial records. It is found that subsequence aftershocks increased the ductility demands on the EPP structures [2]. Similar conclusion was drawn when researchers extended the studies to multi-degree-of-freedom (MDOF) steel frame buildings using repeated earthquake ground motions [3]. In the aforementioned studies, aftershocks were generated based on ground motion characteristics of main shock assuming such characteristics were shared between main shock and aftershocks. However, recent researches had found that ground motion characteristics, such as predominant period, duration and frequency content, of main shock and aftershocks are different for all seismic scenarios [4]. Therefore, this numerical study will focus on investigating the seismic performance of SMA-retrofitted bridge when subject to a sequence of real main shock and aftershock ground motions. The change of demand at bridge s components such as expansion joints, bearings, and abutments will be monitored. The sequence of development of plastic hinges among piers will be tracked and retrofitted accordingly. Prototype bridge geometry The bridge structure considered in this study is a 4-span box girder bridge designed according to old bridge seismic design provisions (prior to 1971). Figure 1 shows the geometry of bridge structure and the cross section of bridge superstructure and piers. The bridge piers are labeled with numbers 1 to 4 as shown in Figure 1.a, and will be referred to herein by its pier number. Total span length of bridge is m. The superstructure is standard AASHTO box-girder segment with depth of 1.8 m and width of 12.9 m. All bridge piers are 1.52 m in diameter. The longitudinal reinforcement ratio is 2% and transverse reinforcement is #4 spirals at 305 mm spacing for all piers. Expansion joints with 25.4 mm of gap are placed in between frames of the bridge. Expansion joints are labeled with numbers 1 to 3 as shown in Figure 1.a and will be referred to herein by its joint number. Lead rubber bearings are located at the abutment and expansion joints. All bearings are sized according to Federal Highway Administration (FHWA) Bridge Design for their load carrying capacity and stability [5]. The size of all bearings is 508 mm by 607 mm. Seat-type abutments are considered in this study with width of 12.9 m and typical height of 1.8 m mm of gap is presented between superstructure and the abutments. The single column bents with different dynamic characteristics, due to their varying height and span length, are considered in order to investigate the effects of pounding in bridges under sequential seismic loading. The bents have heights of 6.1 m, 15.2 m, 6.1m, and 12.2 m. Due to the gapped seat-type abutments considered in this study, there is possibility of pounding not only between frames of bridges, but also between the deck and abutment on each end of bridge.

4 Figure 1a. Elevation view of bridge. Figure 1b. Cross-section of pier. Figure 1c. Cross-section of box-girder. Analytical model The analytical model of bridge was developed using a finite element analysis program, OpenSees [6]. The bridge model was composed of linear beam-column elements for superstructure and nonlinear displacement-based beam-column elements for bridge piers. RigidLink element was used to represent the pier cap which connects between the geometric centroid of superstructure and the top of pier. Cross-section of bridge piers was modeled using fiber sections with appropriate material properties being assigned to cover concrete, core concrete, and reinforcing bars. Classical damping was assumed with 2% damping ratio assigned to first mode. All bridge piers were fixed at base. All elements, except superstructure elements, were expected to behave nonlinearly. Material models and cross section OpenSees Concrete04 uniaxial material was used to describe the behavior of both confined and unconfined concrete. In the as-built model, Mander et al. (1988) model of confined concrete was used to determine the ultimate strain of core concrete with internal spiral confinement. For the retrofitted portion of the column, Mander et al. model of confined concrete was modified to account for additional confinement provided by external SMA spirals on both cover and core concrete. Both the active and passive confining pressures of SMA spiral were considered. Same diameter but different spacing of SMA wire was utilized to achieve three levels of confinement. Table 1 presents the diameter and spacing of SMA wires used and the resulting confining pressure acting on core and cover concrete. The effective confining pressure exerting on core concrete ranged from 0.6 MPa to 2.67 MPa. Reinforcing bars were represented using Steel02 uniaxial material model with isotropic strain hardening effect. Yielding stress of MPa was assumed for all reinforcing bars.

5 Table 1. Properties of SMA spirals and confined concrete for the three confinement levels assumed in the study. Confinement Level Low Intermediate High Concrete type Core Cover Core Cover Core Cover Wire Diameter (mm) Spacing (mm) Peak concrete stress, f' cc (MPa) Concrete strain at peak stress, e' cc Ultimate concrete stress, f cu (MPa) Ultimate concrete strain, e' cu Effective confining pressure (MPa) Expansion joints Bi-linear behavior with strain-hardening ratio of 10% was used to represent the forcedeformation relationship of a single lead-rubber bearing. The area of lead presented in leadrubber bearing was assumed to be 10% of total bearing area according to Robinson. The lead was assumed to yield at a shear stress of 10.5 MPa [7]. Based on the area determined by FHWA Bridge Design Example, the resulting initial stiffness and yielding strength for bearings located at expansion joints were 24 kn/mm and 325 kn, respectively. The resulting element was then connected in parallel to a gap element with high linear stiffness to represent the gap presented between the superstructures and to capture the poundings that may occur at the interface. A relatively high linear stiffness was assumed for the gap element to prevent penetration of superstructures during pounding. Bridge-Abutment interaction The force-deformation curve of abutment backfill soil was developed using Hyperbolic Forcedisplacement (HFD) backbone curve proposed by Shamsabadi [8]. The HFD backbone curve represents the passive soil pressure mobilized as abutment moves into the backfill soil. Soil pressure resulting from soil located at back and front of abutments were considered. The difference in soil height was taken into account by changing the height parameter in the HFD method. The backbone curve was then modeled with uniaxial material Hyperbolic Gap material in OpenSees and assigned to a zero-length element. The lead-rubber bearings located at two abutment ends were modeled the same way as bearings at expansion joints. The initial stiffness and yielding stress of bearings at abutment ends were 13 kn/mm and 227 kn, respectively, calculated based on the same assumptions mentioned in previous section. The zero-length springs representing bearings and abutment response were placed in series and connected to a fixed end.

6 Ground motions and analysis of the seismic response Total of three sets of real main shock and aftershock strong ground motions were considered. Table 2 presents important characteristics of the selected ground motion records. The acceleration response spectra of the selected ground motion records are presented in Figure 2. In this study, three limits states were considered for the bridge model. The limit states are: 1) crushing of core concrete, 2) ultimate capacity of abutment, and 3) hinge opening exceeding 305 mm. The first limit state indicated the development of plastic hinge at a pier. The second limit state was adopted from Caltrans Seismic Design Criteria (SDC) which defined the ultimate longitudinal force capacity of an abutment as the effective area of abutment multiplied by the maximum passive pressure of soil. The height proportionality factor was then applied to account for different height of abutment [9]. For this study, the ultimate capacity of abutment was determined as 5653 kn. Lastly, short seat widths, 152 mm to 305 mm typically, were provided at expansion joints for bridge designed in compliance to older design standard [10]. In this study, maximum hinge opening of 305 mm was chosen as the last limit state of the bridge to prevent the unseating of superstructure. Table 2. Characteristics of selected ground motion records. Earthquake Chi-Chi (1999) Christchurch (2011) Irpinia (1980) Ground motion type PGA (g) Predominant Period (s) Housner Intensity (mm) Significant Duration (s) Main shock Aftershock Main shock Aftershock Main shock Aftershock Spectrum Acceleration (g) Period (sec) Figure 2. Spectrum accelerations of selected ground motion records. Christchurch - 01 Christchurch - 02 Chi-Chi - 01 Chi-Chi - 02 Irpinia - 01 Irpinia - 02 In order to quantify the effectiveness of SMA retrofitting and to investigate the effect of frequency content of ground motion on structural response of RC bridge, Incremental Dynamic Analysis (IDA) was used in this study [11]. The intensity parameter used to scale the ground motion records was PGA. Based on the framework of IDA, an analysis procedure was developed

7 to investigate the change of critical component of bridge after one or more piers of the bridge were retrofitted. The analysis procedure started with determining the maximum intensity of each given main shock that can be resisted by the as-built bridge model without reaching any of the limit states. This was accomplished by subjecting the bridge to the main shock with incrementally increasing intensity until the bridge reached any of the limit states. The maximum intensity (PGA max ) of main shock was then reduced by 20% as a factor of safety to ensure that some amount of bridge capacity was reserved when the subsequent aftershock ground motion was applied. A sequence of ground motions consisting of main shock and aftershock was considered in second stage of the analysis procedure. The intensity of main shock (i.e. 0.8 x PGA max ) was held constant while the intensity of aftershock increased incrementally until any of the limit states was reached. If the crushing of core concrete occurred, the damaged part of pier (regardless whether retrofitted or not) would be retrofitted with next level of confinement. A new analysis with same intensity of main shock and larger intensity of aftershock was then carried out on the undamaged and retrofitted bridge model. By implementing such analysis procedure, the sequence of damage under increasing intensity of aftershock was monitored and retrofitted accordingly. The analysis procedure continued until one of the following happened: 1) core concrete crushing occurred at a location where high level of confinement was applied, 2) capacity of abutment was reached, or 3) hinge opening exceeding 305 mm. At that point, the bridge model would be considered to reach its failure point. The same analysis procedure was repeated for each of selected earthquakes records. Results and discussion Based on the first stage of analysis procedure, it was found that the maximum PGA (PGA max ) of main shock which the as-built bridge model can resist were 0.42g, 0.73g, and 0.61g for Chi-Chi, Christchurch, and Irpinia earthquakes, respectively. In the second stage of analysis, a sequence of main shock with constant intensity (0.8 PGA max ) and aftershock with increasing intensity was applied. Figure 3 illustrates the SMA confinement level that was applied to all four bridge piers at various PGA of aftershock when the bridge was subjected to Chi-Chi, Christchurch, and Irpinia earthquakes. In the above mentioned figure, the lowest values shown on the scale for aftershock PGA correspond to the PGA of aftershock at which the onset of core concrete crushing first occurred at any location of as-built pier. Therefore, these aftershock PGAs indicate the maximum intensity of aftershock the bridge can withstand without any SMA confinement. The results show that crushing of core concrete occur first at piers supporting frames with the lower fundamental periods. Even though Frames 1 and 3 had similar dynamic characteristics due to their identical span length and height, the interaction with adjacent frames and abutments caused the pier supporting Frame 3 to be more vulnerable than the pier supporting Frame 1. Frame 3 was adjacent to two relatively flexible frames (Frames 2 and 4) while Frame 1was located next to the abutment. Therefore, the drift demands on Frame 3 were higher compared to Frame 1.

8 PGA for aftershock (g) Pier Number Figure 3a. Chi-Chi earthquake. PGA for aftershock (g) Pier Number Figure 3b. Christchurch earthquake PGA for aftershock (g) Pier Number Figure 3c. Irpinia earthquake. Figure 3. SMA confinement applied to bridge piers at various PGA of aftershock The Irpinia ground motions amplified uniformly across a wide range of fundamental periods as shown in Figure 2, therefore, the level of retrofitting needed for Piers 1 and 3 were identical for aftershock PGA of 0.82g or lower. With aftershock PGA higher than 0.82g, the level of confinement needed for Pier 3 was consistently higher than that for Pier 1. This result illustrates that a combination of fundamental periods of frames, interactions with adjacent frame and characteristics of ground motions all contributed to the level of confinement needed for bridge piers. As mentioned in previous section, the lowest values shown on scale for aftershock PGA in Figure 3 corresponded to maximum intensity of aftershock the bridge can resist without any SMA confinement. Hence, any increases of aftershock PGA the bridge can resist represent the additional capacity of the bridge resulting from SMA confinement. The improvement of bridge s overall performance can then be quantified as the ratio between increased aftershock PGA and the maximum aftershock PGA the as-built bridge can resist. Improvements in bridge s overall performance of 88%, 117% and 71% were achieved at failure point of the bridge, which was defined as onset of crushing of core concrete at location where high level confinement was applied, for Chi-Chi, Christchurch, and Irpinia earthquakes, respectively. Significant improvements, namely 50%, 87% and 42% for Chi-Chi, Christchurch, and Irpinia earthquake,

9 were obtained with intermediate SMA confinement provided to Piers 1 and 3 and no SMA confinement provided to Piers 2 and 4. The reserved capacity of bridge was limited once plastic hinge was formed at Pier 4. After Pier 4 was retrofitted, there were only 38%, 30%, and 29.7% increases in performance of the bridge for all above mentioned earthquakes. For the three earthquakes considered, no SMA confinement was needed for Pier 2 until the failure point was reached. This demonstrates that the overall seismic performance of the bridge is dominated by its most vulnerable pier. The results also indicated the necessary SMA confinement needed for each of the four piers to achieve a given target performance of the bridge. Based on results from all three earthquakes, one can easily realizes that only low SMA confinement is needed for Pier 3 if the target performance of the bridge is to resist an aftershock with PGA of 0.5g or lower. If the target performance is to resist an aftershock with PGA of 0.6g or lower, then minimum of low and intermediate SMA confinements are needed for Piers 1 and 3, respectively. The results provide insightful knowledge that allow important engineering decisions to be made on level of retrofit needed for the bridge to achieve a given target performance. Hinge opening and pounding force As intensity of aftershock increases, the hinge openings at all three expansion joints increase. The maximum hinge opening occurs at Joint 1, between Frame 1 and Frame 2, where the difference of fundamental periods is the greatest. The maximum hinge opening at the most critical joint, Joint 1, and maximum abutment displacement at various aftershock PGA are shown in Table 3. The maximum hinge opening of mm occurred when the bridge was subjected to Irpinia earthquake with aftershock PGA of 0.99g. The time history of hinge openings for all three expansion joints are shown in Figure 4. The maximum hinge opening did not exceed 305 mm, which was the prescribed limit state on relative displacement between frames, and hence, no unseating of superstructure occurred prior to the crushing of core concrete occurred again at location with high SMA confinement. Hinge Opening (mm) Joint 1 Joint 2 Joint Time (sec) Figure 4. Time history of hinge openings during aftershock of Irpinia with PGA of 0.99g. Pounding occurs when the relative displacement, or the hinge opening, between frames and between superstructure and abutment is less than 25.4 mm. As shown in Figure 4,

10 considerable numbers of poundings occur between all frames. For superstructure and abutment interfaces, longitudinal displacement induced at right abutment was higher than that induced at left abutment. This is mainly attributed to the higher drift ratio experienced by the flexible Frame 4, which is located next to the right abutment. Maximum pounding induced at abutments was 5293 kn, which was 94% of the ultimate capacity of abutment. Even through no abutment failure occurred according to the limit state defined in this study, the results have shown that longitudinal displacement demand has been imposed on abutment as the vulnerable bridge piers been retrofitted. Table 3. Maximum hinge openings for Joint 1 and maximum abutment displacement Aftershock PGA (g) Max. Hinge Opening at Joint 1 (mm) Max. abutment displacement (mm) Chi-Chi Christchurch Irpinia Once any of limit states for abutment and expansion joints is reached, there is no further improvement on bridge s overall performance can be achieved through applying external SMA confinement. Hence, the optimal design of retrofit for the bridge is the state of retrofit for all four piers right before any of limit state of expansion joints and abutments is reached. For all three earthquakes considered, optimal design of retrofit is with intermediate SMA confinement is applied to Pier 4 and with high SMA confinement is applied to Piers 1 and 3. Conclusion In this numerical study, the sequence of formation of plastic hinges at different location of RC bridge piers when subjected to strong main shock and aftershock ground motions were monitored and retrofitted accordingly through form of SMA wires. The effect of retrofitted

11 bridge piers on other components of bridge such as expansion joints and abutments were investigated. The results show that as high as 117% of improvement in bridge s overall performance can be achieved through applying SMA confinement. The improvements in performance are limited, namely 38%, 30%, and 29.7% for Chi-Chi, Christchurch, and Irpinia earthquake, after the more flexible Frame 4 is retrofitted. Pier 3 with low confinement has been determined as the level of retrofitting needed for the bridge to resist an aftershock PGA of 0.5g or lower. For a target performance of resisting an aftershock with PGA of 0.6g or lower, the level of retrofitting needed is low and intermediate confinement applied to Pier 1 and Pier 3, respectively. The optimal design of retrofit is identified as Pier 4 retrofitted with intermediate SMA confinement and Piers 1 and 3 retrofitted with high SMA confinement. The hinge opening at expansion joints and longitudinal displacement demand at abutments can be significant as the intensity of aftershock increased. The demand on right abutment reaches as high as 94% of its ultimate capacity. Acknowledgments The authors acknowledge the financial support provided for this research from the National Science Foundation through its Faculty Early Career Development (CAREER) program under Award No Reference 1. Shin, M., & Andrawes, B. Emergency repair of severely damaged reinforced concrete columns using active confinement with shape memory alloys. Smart Materials and Structures 2011; 20(6) 2. Mahin, S. A. Effects of duration and aftershocks on inelastic design earthquakes. Proceedings of the 7th World Conference on Earthquake Engineering 1980; Istanbul, Turk. 3. Fragiacomo, M., Amadio, C., & Macorini, L. Seismic response of steel frames under repeated earthquake ground motions. Engineering Structures 2004; 26(13): Ruiz-García, J. Mainshock-aftershock ground motion features and their influence in building's seismic response. Journal of Earthquake Engineering 2012; 16(5): Federal Highway Administration / National Highway Institutite. LFRD Design Example for Steel Girder Superstructure Bridge (FHWA NHI-04-04) 2003; Washington, D.C. 6. Fenves, G.L., McKenna, F. Scott, M.H., Takahashi, Y. An object-oriented software environment for collaborative network simulation. Proceedings, 13th World Conference on Earthquake Engineering 2004; Vancouver,Canada. 7. Robinson, W. H. Lead-rubber hysteretic bearings suitable for protecting structures during earthquakes. Earthquake Engng. & Struct. Dyn 1982; 10(4): Shamsabadi, A., Khalili-Tehrani, P., Stewart, J. P., & Taciroglu, E. Validated simulation models for lateral response of bridge abutments with typical backfills. Journal of Bridge Engineering 2010; 15(3): California Department of Transportation (CALTRANS). Seismic design criteria (SDC Version 1.6) 2010; Redwood, California. 10. California Department of Transportation. Visual Catalog of Reinforced Concrete Bridge Damage 2006; Redwood, California. 11. Vamvatsikos D, Allin Cornell C. Incremental dynamic analysis. Earthquake Engineering and Structural Dynamics 2002;31(3):