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Engineering Structures 32 (2010) 21922209

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Engineering Structures

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Seismic assessment of an existing non-seismically designed majorbridge-abutmentfoundation system

Aman Mwafy a, Oh-Sung Kwon b,, Amr Elnashai c

a Department of Civil and Environmental Engineering, United Arab Emirates University, P.O. Box. 17555, Al-Ain, United Arab Emiratesb Department of Civil, Architectural and Environmental Engineering, Missouri University of Science and Technology, Rolla, MO 65409, USAc Department of Civil and Environmental Engineering, University of Illinois at Urbana-Champaign, Urbana, IL 61801, USA

a r t i c l e i n f o

Article history:

Received 14 May 2009

Received in revised form

13 March 2010Accepted 15 March 2010

Available online 18 April 2010

Keywords:

Major highway bridges

Seismic assessment

Inelastic dynamic analysis

Soilstructure interaction

High seismicity regions

Bridge deficiencies

a b s t r a c t

A comprehensive study carried out to assess the seismic response of a 59-span bridge using a refinedinelastic modeling approach and considering SoilStructure Interaction (SSI) is summarized in this paper.Thefocus is on describing themethodology adopted to idealizethe bridge andits foundationsystem,whileonly highlights from the extensive elastic and inelastic analyses are presented. The bridge represents atypical case of vulnerable complex bridges since it was built in the early seventies with minimal seismicdesign requirements at a distance of about 5 km from a major fault. The SSI analysis is significant inthis study due to the length of the bridge, the massive and stiff foundation, and the relatively soft deepsoil of the site. A series of three-dimensional dynamic response simulations of the entire bridge areconducted using several analysis tools to verify the developed analytical models. The performance-basedassessment study employs 144 site-specific input ground motions representing three seismic scenarios,corresponding to 500, 1000 and 2500 years return periods, to identify areas of vulnerability in the2164-meter bridge at various hazard levels. It is concluded that the seismic response of the bridge atthe 500 years ground motions does not meet todays standards, while the demands under the effect ofthe 1000 years ground motions almost exceed the capacity of most bridge components. The demandssignificantly increase under the effect of the 2500 years earthquake scenario and considerably exceedthe collapse limit states. The results clearly reflect the benefit of retrofitting different bridge componentsto mitigate the anticipated seismic risk. The presented assessment study contributes to improve publicsafety by exploiting the most recent research outcomes in predicting the seismic response of complexhighway bridges, which are essential for developing reliable and cost-effective retrofit strategies.

2010 Elsevier Ltd. All rights reserved.

1. Introduction

Research carried out during the past two decades has led tosignificant changes in seismic design provisions of bridges. Theintroduction of the AASHTO Load and Resistance Factor Design(LRFD) Specifications [1] and the Guide Specifications for LRFD

Seismic Bridge Design [2] is aimed at providing more uniformsafety for different types of bridge systems. The release of theFHWA Retrofitting Manual for Highway Structures [3] also pro-vides comprehensive procedures for assessment and retrofittinghighway bridges based on recent experiences in the US, Japan,and other countries. This major revision in the FHWA SeismicRetrofitting Manual introduces a performance-based retrofit phi-losophy and defines several methods for the detailed evalua-tion of existing bridges. These revisions in design specifications

Corresponding author. Tel.: +1 573 341 4536; fax: +1 573 341 4729.

E-mail address: [email protected] (O.-S. Kwon).

and retrofitting guidelines draw attention to the need for seis-mic assessment of complex highway bridges designed to preced-ing provisions to determine the level of risk associated with loss ofserviceability or possible damage. This is particularly significant inthe light of the continuous updates in seismic hazardmaps for sev-eral regions (e.g. [4]).

A large number of bridges were designed and constructed ata time when bridge codes were insufficient according to currentstandards. The deficiencies in highway bridges designed prior tothe seventies result in excessive seismic displacements and largeforce demands that were substantially underestimated. The antic-ipated damage includes unseating and pounding of superstructureat abutments and expansion joints, shear and flexural failure ofRC piers, beamcolumn joint failure, footing and abutment failureand amplification of response due to soft/liquefiable soil (e.g. [5]).The existing bridge inventory designed to previous provisions isthus likely to suffer damage when subjected to seismic scenarioscomparable to those observed in severe earthquakes (e.g. San Fer-nando,USA, 1971; Loma Prieta, USA, 1989;Northridge, USA, 1994).

0141-0296/$ see front matter 2010 Elsevier Ltd. All rights reserved.doi:10.1016/j.engstruct.2010.03.022
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A. Mwafy et al. / Engineering Structures 32 (2010) 21922209 2193

In-depth performance-based seismic assessment studies of ma-jor highway bridges point the way towards improving the under-standing of the complex seismic performance of similar structures.This performance-based evaluation approach requires bridges tosatisfy different performance criteria for different levels of groundmotion. For instance, the bridge may suffer minor damage butshould be operational under frequent earthquakes with low inten-sity. Under infrequent earthquakes with large intensity, the bridgeshould provide an acceptable level of life-safety. Quantifying thelevel of risk associated with anticipated earthquake scenarios en-ables taking rational decisionsto retrofit,replace or accept the risk.

The preliminary screening and prioritization approach propo-sed by the FHWA Retrofitting Manual for Highway Structures [3]classifies bridges according to their Seismic Retrofit Categories(SRC) to identify minimum requirements for screening, evaluationand retrofitting. The FHWA Retrofitting Manual recommends per-formance criteria according to bridge importance and anticipatedservice life, with more rigorous performance required for impor-tant, relatively new bridges, than for standard bridges at the end oftheir service life. The manual also introduces a performance-basedretrofit philosophy similar to that used for the performance-baseddesign of new buildings and bridges. Following the preliminary

screening process, the FHWA retrofit manual [3] recommends sixevaluation procedures, with increasing order of rigor, for quan-titatively assessing the seismic performance of existing highwaybridges. Method E, which is based on the inelastic time historyanalysis, compares seismic demands with member capacities anddetermines the degree of damage to the structure at the desiredearthquake levels. The latter method, which requires considerablecomputational effort and a significant level of skill to interpret theresults, is recommended for irregular complex bridges with thepotential for substantial inelastic behavior particularly when site-specific ground motions are used. The major bridge assessed in thepresent studyrequires the most rigorous evaluation procedure rec-ommended by the FHWA retrofitting manual. The primary objec-tives of this paper are therefore as follows:

Discuss the comprehensive methodology adopted to realisti-cally assess the seismic response of a major bridge and its foun-dation system taking into account the most recent advancesin the analytical modeling approaches of substructure, super-structure, abutments and foundations.

Present sample results from the comprehensive vulnerabilitystudy to highlight the significance of advanced simulationapproaches in identifying areas of vulnerability of complexbridges.

Investigate the significance of the recent understanding aboutthe seismicity of the site of the investigated bridge in providinginsight into its seismic response at different hazard levels.

2. Description of the I-155 bridge at Caruthersville

Bridge A-1700 at Caruthersville, which carries Route I-155across the Mississippi River between Pemiscot County, Missouri,and Dyer County, Tennessee, is a major bridge that has a press-ing need for detailed vulnerability assessment [6]. Although the59-span 2164-meter bridge is about 5 km from a presumed ma-

jor fault, it was constructed in the early seventies with minimalseismic design requirements by todays standards. According tothe design information [7], Peak Ground Accelerations (PGA) of0.10gand 0.06gwere employed in the seismic design of thebridgein the transverse and longitudinal directions, respectively. Clearlythis reflects the limited knowledge about the seismicity of the con-struction site at the time of designing the bridge; the whole fieldof study has advanced since the time of construction. The super-

structure consists of eleven units separated by expansion jointsand supported on a variety of elastomeric and steel bearings. The

main channel crossing is composed of two-span asymmetrical can-tilever steel truss and ten-span steel girders, while approach spansare precast prestressed concrete girders. The substructure includespiers on deep caissons and bents on steel friction piles driven intothenear surface silty sands andclayey materials. Bedrock is locatedmore than 800m below the sand, gravel, andhard clay strata. Fig. 1shows a three-dimensional view of the I-155 bridge and pictori-ally summarizes the adopted three-dimensional analytical mod-eling approach, which is discussed in the following sections. Thisbrief description of the Caruthersville Bridge highlights the press-ing need to reliably assess its response under anticipated seismicscenarios.

3. Development of simulation models for inelastic analysis

Detailed three-dimensional dynamic response simulations ofthe entire bridge including foundations and soil effects arecarried out using a number of verified analysis platforms. Thefinite element analysis programs SAP2000 [8] and ZEUS-NL [9]are employed for elastic and inelastic analysis of the structure,respectively. The Pacific Earthquake Engineering Research (PEER)Center analysis platform OpenSees [10] is used for an inelastic

simulation of the foundation and the underlying sub-strata. TheSAP2000 analytical models are mainly employed for verificationsof the ZEUS-NL fiber model before executing the extensiveinelastic analysis. ZEUS-NL is mainly employed to estimate thecapacities and demands from inelastic pushover and responsehistory analyses. The latter finite element analysis platform wasdeveloped at Imperial College London and at University Illinois atUrbanaChampaign, and has been extensively employed in largeprojects (e.g. [11,12]).

3.1. Super- and sub-structure modeling

Three different analytical models are developed for the bridge:(i) SAP2000 detailed model, (ii) SAP2000 simplified model and

(iii) ZEUS-NL fiber model. The first model is developed to repre-sent all sub- and super-structural components for elastic analy-sis. Steel and concrete cross-sections from the SAP2000 library areemployed to realistically model different structural members forelastic analysis. This modeling approach, particularly for the su-perstructure, is computationally demanding for inelastic responsehistory analysis. Also, the design philosophy of bridges relies onbridge piers to dissipate energy rather than the superstructure,which remains elastic. The detailed SAP2000 analytical model istherefore modified to reduce the number of elements and nodesto a manageable limit for inelastic analysis. With the exception ofspans with truss, the superstructure is replaced with a number ofcross sections with equivalent geometrical properties connectedtogether using rigid arms. This simplification in the superstructure

enabled reducing the number of elements and DOFs by about 50%.On the other hand, substructure members are refined by subdi-viding the columns to a number of elements to accurately moni-tor the inelastic response during timehistory analysis. Moreover,the SAP2000 joint constraints, which are not available in ZEUS-NL,are replaced with strong arms. The simplified SAP2000 model wastransferred to ZEUS-NL for inelastic analysis. Due to the complexbehavior of the truss, it was transferred to ZEUS-NL without sim-plifications.

Based on available cross-sections in the library of SAP2000 andZEUS-NL, equivalent cross-sections are adopted for modeling of thebridge members. In the detailed ZEUS-NL model, each structuralmember is assembled using a number of cubic elasto-plasticelements capable of representing the spread of inelasticity within

the member cross-section and along the member length via thefiber analysis approach. Sections are discretized to steel, confined


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2194 A. Mwafy et al. / Engineering Structures 32 (2010) 21922209

b

a

c d

Fig. 1. Three-dimensional analytical model of the I-155 bridge at Caruthersville, Missouri, including foundation and abutment effects: (a) three-dimensional simulation

model of the entire bridge; (b) components of the main channel crossing; (c) OpenSees abutment model and sample pushover analysis results; (d) OpenSees model and

sample pushover results of one of the bridge footings.

and unconfined concrete fibers. The stressstrain response at each

fiber is monitored during the entire multi-step analysis. Therefore,the degradation of stiffness with increasing inelasticity is fullyaccounted for in the ZEUS-NL inelastic analysis.

Equivalent gravity loads and mass are distributed in the ZEUS-

NL fiber model on the superstructure and along the piers height.The employed distributed mass elements utilize cubic shape func-tion and account for both the translational and rotational inertia.

The total weight of the bridge is 1562,549 kN (351,275 kip), whichincludes superstructure, substructure, non-structural members,pile caps andcaissons. The corresponding mass is 159,336 kN s2/m(10,918 kip s2/ft). The superstructure weight is higher than the

substructure in the approach spans, which is not the case in thesteel girders and the truss spans. This is due to the lower weight ofthe steel members compared with the concrete counterparts and

the lightweight deck of the steel and truss spans. Also, the mas-sive weight of the caissons significantly increases the substructureweight of the steel and truss spans.

As a result of the several deficiencies observed in structural

members in the latest available inspection report of the bridgeand the lack of reliable information confirming the actual materialcharacteristics, nominal material properties are used in analysis.

The cylinder concrete strength (f

c) is 21 MPa (3000 psi) for nor-mal concrete and 28 MPa (4000 psi) for prestressed concrete. The

yield strength of reinforcing steel (fy) is 40,000 psi. Structural car-bon steel ASTM A36-69 (fy = 248 MPa or 36,000 psi) is used fortruss members and steel beams. A bilinear model was used to ide-alize steel members and reinforcement. In this model, the loadingandunloading in theelastic range followa linearfunction through-out various loading stages with constant stiffness represented bythe Youngs modulus of steel. In the post-elastic range, a kinematichardening rule forthe yield surface defined by a linear relationshipis assumed [13]. A uniaxial constant active confinement concrete

model was employed in the ZEUS-NL analytical model. This modelhas a good balance between simplicity and accuracy and includesenhancements in the cyclic degradation rules, inelastic strain, andshape of unloading branches [14]. The model, which incorporatesthe influence of confinement effects on the peak stress and strainas well as on thepost-peak stressstrain relationship, canprovide agood estimation of the cyclic response of RC members under cyclicand dynamic loading.

As suggested from the bridge drawings, appropriate transla-tional () and rotational (r) Degrees of Freedom (DOFs) are re-leasedat thewind transfer deviceof thetruss, as shown from Figs.1and 2. The translational DOFs, X, and the rotational DOFs, rY andrZ, of the stringers and the deck are also released at certain ex-pansion joints as indicated in the bridge drawings. Bridge bear-

ings and expansion joints are realistically modeled using ZEUS-NLjoint elements. The rotation about the transverse axis, rY, is only


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(a) Released DOFs at the truss intermediate hinge. (b) Gap elements controlling displacement.

Fig. 2. Modeling approach of the truss wind transfer device.

'' ''

(a) ZEUS-NL model of the expansion shoe at Bents

15, 19 (left), 21 (right) and 25.

(b) ZEUS-NL model of the Bronze self-lubricating

bearings at Bents 8, 14 (left), 26 (right), 31, 41 and51.

Fig. 3. ZEUS-NL models of bridge bearings.

allowed at the fixed shoes. Based on the drawings, these bear-ings are located at the substructuresuperstructure connection ofBents 27, 913, 14 (right), 1618, 20, 21 (left), 2224, 26 (left),2730, 3240, 4250 and 5259. On the other hand, the slidingbearings are located at the two abutments and at the substruc-turesuperstructure connection of Bents 8, 14 (left), 15, 19, 21(right), 25, 26 (right), 31, 41 and 51.

Modeling of the steel expansion shoes at Bents 15, 19 (left), 21(right) and 25 follows the analytical model suggested by Manderetal. [15]. This bearing type is highlysusceptibleto instability since

the horizontal motion in the longitudinal direction is accommo-dated through a rocking motion. This motioncausesfrictionresult-ing from a rolling resistance at the base and a Coulomb friction atthe top hinge [16,17]. Mander et al. [15] observed that debris anduneven wear cause the experimental hysteresis loops to be irreg-ular. The ZEUS-NL model of this bearing is shown in Fig. 3(a). Theinitial stiffness of this bearing, K1, is 14 kN/mm with a post-yieldstiffness, K2, of 0.018k1. The frictional coefficient is 0.04. The ana-lytical model of the expansion bearing at Pier 19 (right) has also aninitial stiffness, K1,of14kN/mm with a perfectly plastic post-yieldstiffness. However, the longitudinal displacement of this bearingis restricted to 6.0 in. The frictional coefficient is 0.20, as recom-mendedfor the low-type expansion bearings investigatedby Man-der et al. [15].

Movable bearings, which typically have small friction coeffi-cient at the low velocity rates, may have higher friction under high

seismic deformation. For un-lubricated elastomeric bearings, thiscoefficient at high velocities ranges from 5% to 15%, or even higherat the low temperature (e.g. [18]). It was also concluded in pre-vious experimental studies that the coefficient of friction slightlydecreases again under the high velocities due to frictional heating.The Bronze Self-lubricating bearings at Bents 8, 14 (left), 26 (right),31, 41 and 51 are idealized using the model shown in Fig. 3(b). Theinitial stiffness, K1, is taken equal to 123 kN/mm based on the lowtype sliding bearings tested by Mander et al. [15]. The frictionalcoefficient of the Bronze Self-lubricating surface is taken equal to

0.10. The post yield stiffness is estimated from the geometry of thesystem. The horizontal movement of this bearing implies a verti-cal uplift of the superstructure. The post yield stiffness, K2 = W/r,where W is the weight and r is the radius of the bearing surfacecurvature. The top plates of these bearings are slotted to allow for3.5 in. of movement. The displacement in the longitudinal direc-tion is therefore restricted to this limit and the stiffness of the sys-tem significantly increases (K3) if the maximum displacement isreached, as shown from Fig. 3(a).

The structural gaps at the abutments and the expansion jointsare considered in the inelastic analysis performed using ZEUS-NL. In the latter modeling approach, joint elements with tri-linearasymmetric elasto-plastic idealization capable of representing theslippage and collision are employed. Two modeling approaches

were investigated to idealize the impact of stiffness at the expan-sion joints. In the first approach, the pounding is represented by a


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Fig. 4. ZEUS-NL model of the expansion joint at Abutment 1, Bents 8, 14 (left), 26 (right), 31, 41, 51 and Abutment 60.

Fig. 5. Vertically exaggerated view of soil profile along the length of the bridge.

nonlinear spring (Hertz model). The impact stiffness increases inthis approach from 2.92E6 MPa to 4.38E6 MPa at a penetration of1.27 mm, and increases again at 2.54 mm to 8.76E6 MPa (e.g. [ 19,20]). These values are controlled to ensure that the penetration ofpounding is less than 2.54 mm (0.1 in.). Results obtained from thisidealization were compared with a more simplified linear springmodel employing the maximum impact stiffness of 8.76E6 MPa. Itwas concluded that the former model has a marginal effect on theresponse compared with the simplified linear model. It was there-fore decided to use the simplified approach to model the impactstiffness. Fig. 4 shows the force versus relative displacement rela-tionships of the joint elements representing the expansion joint. Inthis modeling, a positive relative displacement corresponds to anopening of the joint gap and a negative displacement correspondsto a closing of the gap. When the gap at the abutment and at theexpansion joint undergoes a relative movement in the negative di-rection (joint close) exceeding the gap width, the joint element be-gins resisting further opening (collisions).

The modes of vibration obtained from modal analysis reflectthe high contribution of the steel truss to the seismic responseof the bridge, as subsequently discussed. Accordingly, 4% dampingratio is considered in the SAP2000 model used for elastic analysis.The mass-proportional and stiffness-proportional parameters arecalculated for Response History Analysis (RHA) based on thepredominant periods of the structure in the two principledirections.

The damping and material modeling are critical components inthe prediction of inelastic seismic structural response [21]. Hence,they were thoroughly investigated before executing the ZEUS-NLinelastic analysis. Hysteretic damping, which is responsible for thedissipation of the majority of energy introduced by the earthquakeaction, is accounted for in the inelastic fiber formulation of the

inelastic elements. A relatively small quantity of non-hysteretictype of damping should be also added to the inelastic model.

This is achieved through stiffness-proportional damping [22].1.0%, 1.5% and 2.0% stiffness-proportional damping ratios areinvestigated. It is noteworthy that these damping ratios wereapplied on substructure members, which are the main sourceof energy dissipation. A higher level of damping is applied onthe superstructure, which is anticipated to remain in the elasticrange. To compare the impact of damping on the response ofthe Caruthersville Bridge obtained from SAP2000 and ZEUS-NLinelastic analysis, an input ground motion developed to matchthe Uniform Hazard Spectrum (UHS) of the bridge for a 500 yearsreturn period is used. The record was scaled to a Peak GroundAcceleration (PGA) of 0.05g to avoid a significant inelasticityfrom the ZEUS-NL nonlinear analysis. The objective from thecomparison is to tune the ZEUS-NL damping level and to verifyits results with those obtained from the SAP2000 elastic analysis.The results indicated that the ZEUS-NL results are comparable tothose obtained from SAP2000 when a 2% damping level is used.The displacement andthe base shear demands from ZEUS-NL were

significantly higher than those from SAP2000 when a lower levelof damping is used. Based on these comparisons, it was decided toemploy a 2% stiffness-proportional damping in inelastic analysis.This damping ratio is applied on substructure, which is the mainsource of energy dissipation. Higher level of damping is applied onsuperstructure, which is anticipated to remain in the elastic range.

3.2. Soil and foundation modeling

Soil and foundation may have significant impact on the seismicresponse of bridge structures, particularly those with stiff founda-tion and relatively soft deep soil [23]. Refined inelastic simulationsof the foundation and the underlying sub-strata are undertakenusing the Pacific Earthquake Engineering Research Center analy-

sis platform OpenSees [10]. The objective is to realistically esti-mate thesoil properties required forSAP2000 andZEUS-NLsoil and


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zy

x

zy

x

z

yx

Fig. 6. Different bridge footing profiles.

foundation modeling. Fig. 5 shows a vertically exaggerated viewof soil profile along the length of the bridge. The site consists of8 types of soil materials. Stratum 1 and 2 are cohesive materialwith unconfined compressive strength of 67 and 48 kPa, respec-tively. This range of strength is for medium to stiff clay. The up-per layers of approach spans are covered by either stratum 1 or 2.These strata are modeled with pressure-independent soil materialmodel. Strata3 to 8 are cohesionlessmaterials. The friction angle of

the materials ranges from 30 to 42. The in-situ properties of thecohesionless material at deeper layer are expected to have higherstiffness due to the large confinement. Since the shear modulus ofeach stratum is given as a constant value regardless of the depthof the stratum, it is assumed that the shear modulus is calculatedat the mid-depth of each stratum. The effective reference pressureat the mid-depth is approximately calculated and used as the in-put parameters for soil and foundation modeling. Strata 3 to 8 arethus modeled with pressure-dependent material, which can dilateor contract with shear deformation depending on the initial den-sity ratio. Based on the density ratio of various soil profiles, strata3and 4 are assumed to be medium sand, strata 5 and 6 as medium-dense sand and strata 7 and 8 as dense sand. The dilatation andcontraction parameters in the pressure-dependent material mod-

els are chosen from the suggested values by the developer of thematerial model [24].

The resistance of soil medium surrounding pile largely dependson the contact surface area normal to the direction of the pilemovement. Due to the large number of piles and the amount ofanalyses required, it is computationally demanding to model eachpile using several finite elements. A single brick element with anequivalent projected area is therefore used for idealizing the pilesection and reducing the computational demands. The equivalentelement used in analysis has the similar properties of the actualpiles in terms of flexural and axial rigidity. Based on the soilprofile, number of piles and batter angle, thirteen soilfoundationprofiles were idealized using OpenSees. Unlike the pile caps atexpansion joint bents, those at other locations have battered piles.

The number of piles varies from 9 to 112, depending on thesupporting loads. Bent 19, 20, and 21 are supported on massive

Fig. 7. Controlling different foundation systems using a single node for inelastic

pushover analysis.

caisson. Fig. 6 describes different bridge foundation profiles. Thefoundation and soil medium are all modeled with 8 node brick

elements.The side boundary of the soil mediumis restrained in thehorizontal translation as shown in Fig. 1(d). Vertical DOFs of sideboundary are released to allow settlement due to gravity loads. AllDOFs of the bottom nodes of the soil medium are restrained.

To reduce the controlling nodes within the pile cap and toprevent local deformations, it is assumed that the pile caps andcaissons behave as rigid bodies. All foundation profiles, exceptcaissons, are thus controlled using a single node connected toeight boundary nodes, as shown in Fig. 7. Symmetry is utilizedto reduce the FE mesh and computational demands for certaintypes of foundation profiles (e.g. Class 6). On the other hand, it wasdifficult to control the response of the caissons using a single nodedueto their massive stiffness. Hence, a numberof nodes areused tocontrol their response, as shown in Fig. 8. Displacement-controlled

pushover analyses are carried out using the above-mentionedrefined FE models to evaluate the loaddeformation relationship of


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ab

Fig. 8. Modeling of the bridge caissons for OpenSees pushover analysis: (a) translational force; (b) rotational moment.

a b

Fig. 9. Idealizing bridge footings using tri-linear relationships: (a) response of footing Class 2 under cyclic and monotonic loading; (b) adopted ZEUS-NL tri-linear model.

each foundation class in Fig. 6. The loaddeformation relationshipsof various foundation profiles are derived from the reactions at thecontrol nodes.

4. Foundation and abutment response under cyclic and mono-

tonic loading

4.1. Bridge foundations

Sample results from the extensive foundationsoil analyses arepresented below. The loaddeformation relationship of foundationClass 2 under the effect of both cyclic and monotonic loading isshown in Fig. 9(a). The thin solid line is from cyclic loading, whilethe thick solid line is from monotonic loading. The results confirmthat the backbone of the hysteretic curve follows the monotonicpushover curve. It was decided based on this comparison toanalyze other foundations profiles under a monotonic loading toestimate their loaddeformation relationships. For linearizationof the nonlinear stiffness of translational DOFs, the maximumforce is assumed to be twice of the reaction from dead loads,which corresponds to inertial force from 2g of horizontalacceleration. Idealized tri-linear relationships are subsequentlyused as soil springs for inelastic analysis. Fig. 9(b) depictsthe adopted ZEUS-NL tri-linear model used to idealize bridgefoundations. Figs. 1012 show the iso-deformation contoursand the corresponding loaddeformation curves obtained fromOpenSees pushover analysis of foundation Classes 5, 6 and 9,respectively. The full results of these comprehensive analysesfor all foundation classes are given elsewhere [25]. Tri-linearidealizations are adopted to simplify the monotonic pushovercurves of different foundation classes. The yield displacement andpost-yield stiffness are chosen so that the tri-linear model closelyrepresents the loaddeformation curve obtained from OpenSeespushover analysis. The idealized loaddeformation relationshipsare therefore used to model the foundation systemin SAP2000 and

ZEUS-NL analytical models. A summary of the foundation stiffnessproperties obtained from the OpenSees analyses is given in Table 1.

4.2. Abutments

The analytical model of the bridge abutments developed usingOpenSees along with the soil profile is described in Figs. 1(c)and 13. Stratum 1 and 2 are cohesive material with undrainedstrength (Su) of 67 and 48 kPa (0.7 and 0.5 tsf), respectively.

All other strata are cohesionless material with friction angleranging from 30 to 42. The inelastic analysis carried out in thelongitudinal and transverse directions of the bridge indicated thatthe response is almost linear. Results of pushover analyses showeda degrading stiffness but at very high level of force. Table 2 showsa summary of the abutment stiffness properties in the longitudinaland transverse direction of the bridge. The vertical translation androtations of the two abutments are fully restrained.

The abutment modeling approach obtained using OpenSees iscompared in Table 3 with a more simplified approach suggestedby Caltrans [26]. In the latter approach, it is assumed that wingwalls do not significantly contribute to the transverse horizontalresistance. Therefore, the transverse horizontal stiffness of theabutment is estimated using the translational stiffness of theabutment pile group. In the longitudinal direction, the abutmentembankment fill stiffness is estimated based on the finding of alarge-scale abutment testing [27]. The initial passive stiffness fromthis testing (11.5 kN/mm/m) is adjusted relative to the back-wallheight [26]. A maximum passive pressure was recommended inthe latter study, allowing the back-wall to break off in order toprevent inelasticity in the foundation system. The stiffness in activeaction is assumed to be one fifth of the passive stiffness [19]. Abilinear elasto-plastic relationship is therefore adopted to modelthe longitudinal behavior of the abutment. The vertical translationand rotations of the abutment are fully restrained. It is clear fromthe comparison shown in Table 3 that the refined OpenSees modelresults in a significantly different response at the investigatedlow level of input ground motion (PGA = 0.05g). This confirmsthe significance of the refined modeling approach adopted in

the present study for different components of the CaruthersvilleBridge.


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Fig. 10. Iso-deformation contours and the corresponding loaddeformation curves obtained from the inelastic pushover analysis of foundation Class 5.

5. Performance criteria and analysis procedures

To assess the seismic performance from inelastic static

and dynamic analysis results, reliable definition of responseparameters is needed. Two particular limit states in the response

of the bridge are required to be defined; one at which significantyield occurs and one at which the first indication of failure isobserved. Local yield is assumed when the strain in the main

longitudinal tensile reinforcement exceeds the yield strain ofsteel. The response from the inelastic timehistory analysis is


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monitored and any occurrence of yielding in the main longitudinalreinforcement is reported. On the other hand, the definition offailure or collapse criteria should represent an accurate or aslightly conservative estimation of structural collapse due to its

significance role. The adopted failure or collapse performancecriteria are as follows:

(i) Drift: The drift limit is defined as the ratio of relativedisplacement between the top and the base of the bridge piers totheir height. Several values for the drift collapse limit state havebeen suggested in the literature and in seismic codes. A drift limit

in excess of 3.0% is considered in the present study as an indicationof a global collapse.


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Table 1

Stiffness properties of selected foundation classes.

Footing DOF k0 d0 k1 d1 k2

Class 02

X 1.0190E+06 9.0620E04 6.6800E+05 4.5310E03 3.9710E+05

Y 8.8900E+05 1.3240E03 4.7530E+05 6.6210E03 2.1890E+05

Z 1.5880E+06 3.4810E04 1.5310E+06 1.7400E03 1.4130E+06

rx 3.8420E+06 1.0000E03 1.7650E+06 6.0000E03 1.0750E+06

ry 5.6010E+06 1.0000E03 3.0330E+06 6.0000E03 1.3120E+06

Class 04

X 1.2460E+06 1.4530E03 7.3990E+05 7.2640E03 3.7110E+05Y 1.1720E+06 1.7910E03 6.0320E+05 8.9570E03 2.6620E+05

Z 2.3180E+06 4.3630E04 2.1950E+06 2.1820E03 1.8170E+06

rx 5.8250E+06 1.0000E03 2.7070E+06 6.0000E03 1.2540E+06

ry 1.2630E+07 1.0000E03 5.6610E+06 6.0000E03 1.5270E+06

Class 05

X 1.4600E+06 9.6050E04 1.0180E+06 4.8030E03 6.0370E+05

Y 1.3990E+06 1.0670E03 9.3220E+05 5.3330E03 5.1470E+05

Z 2.3980E+06 3.6600E04 2.3120E+06 1.8300E03 2.1600E+06

rx 5.5420E+06 1.0000E03 3.0980E+06 6.0000E03 1.6860E+06

ry 1.3330E+07 3.5200E03 7.0020E+06 8.8000E03 3.0680E+06

Class 06

X 4.5970E+06 5.0000E03 2.3350E+06 2.5000E02 1.5070E+06

Y 4.0710E+06 5.0000E03 1.6960E+06 2.5000E02 1.0480E+06

Z 1.0070E+07 2.7800E03 8.4600E+06 1.3900E02 2.4140E+06

rx 1.8470E+08 2.0000E03 8.0370E+07 8.0000E03 3.9390E+07

ry 1.6350E+08 2.0000E03 6.7970E+07 1.0000E02 4.8650E+07

Class 07

X 1.3870E+07 5.0000E03 6.3950E+06 2.5000E02 3.5870E+06

Y 1.2750E+07 5.0000E03 5.7650E+06 2.0000E02 2.4440E+06

Z 1.4590E+07 2.9200E03 8.1440E+06 1.4600E02 1.5810E+06

rx 1.0370E+09 8.8140E04 5.3190E+08 5.2890E03 4.2390E+08

ry 9.9430E+08 8.9480E04 2.8610E+08 4.4740E03 1.6510E+08

Class 08

X 2.3610E+07 8.0000E03 9.1250E+06 4.0000E02 3.8890E+06

Y 2.2670E+07 8.0000E03 8.4070E+06 4.0000E02 2.6990E+06

Z 3.3840E+07 2.5200E03 2.3350E+07 1.2600E02 1.7960E+06

rx 2.6840E+09 5.0000E04 1.7570E+09 3.0000E03 1.3810E+09

ry 3.3160E+09 1.0000E03 8.6050E+08 5.0000E03 3.7480E+08

Class 09

X 1.5830E+07 8.0500E03 5.0920E+06 4.8300E02 1.6560E+06

Y 1.6250E+07 5.9000E03 6.4540E+06 2.3600E02 2.5670E+06

Z 2.3900E+07 4.8650E04 2.1390E+07 2.4320E03 1.8410E+07

rx 1.9730E+09 4.0530E04 1.0300E+09 2.4320E03 8.2870E+08

ry 1.5450E+09 7.1090E04 4.7120E+08 3.5540E03 2.4890E+08

Class 10

X 1.7000E+06 1.4190E03 1.3630E+06 7.0970E03 9.5200E+05

Y 1.5170E+06 1.7630E03 1.1250E+06 8.8160E03 7.1490E+05Z 3.7330E+06 4.8170E04 3.6460E+06 2.4090E03 3.3580E+06

rx 3.2170E+07 1.0000E03 2.1060E+07 6.0000E03 1.3060E+07

ry 8.3320E+07 9.0000E04 5.6540E+07 5.4000E03 3.0010E+07

Unit: force (kN), moment (kN m), length (m), rotation (rad). K0: Initial stiffness. d1: First branch displacement limit. K1: Second branch stiffness. d2: Second branch

displacement limit. K2: Third branch stiffness. Also refer to Fig. 9(b) for the definition of stiffness properties.

Table 2

Summary of abutment stiffness properties.

DOF k0 d0 k1 d1 k2

X 1.3250E+06 2.2180E04 1.2560E+06 1.1090E03 1.0500E+06

Y 9.2060E+05 3.2070E04 8.5390E+05 1.6040E03 7.3660E+05

Unit: force (kN), length (m).

Table 3

Response of the bridge in the transverse direction using different abutment idealization.

Bent Top displacement (mm) Base shear (kN)

Caltrans OpenSees Diff. (%) Caltrans OpenSees Diff. (%)

2 5.2 3.0 74 422 272 55

15 11.7 9.5 23 5 072 4 869 4

20 38.1 22.7 68 42 638 31 153 37

21 23.2 10.4 123 30 666 18 530 65

25 8.1 6.7 22 988 857 15

59 5.9 2.0 195 671 231 191

Response history analysis results from Rec1-500-SE scaled to a PGA of 0.05g.

(ii) Degradation of lateral resistance: Collapse is considered tohave occurred when lateral resistance of a structure drops by morethan 10% below its peak value.

(iii) Formation of hinging mechanism: Such a mechanisminvolves plastic hinges at the base of all bents. Although this

mechanism is difficult to occur, particularly for the long bridgeunder consideration, the spread of yielding in substructuralmembers along the length of the bridge segments, particularly for

the main approach, namely Piers 15 to 21, is considered here as anindication of structural collapse.


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(iv) Capacity versus demand ratios: The capacity of differentbridge components are estimated from extensive pushover

analyses and hand calculations. This is carried out for:

Bridge foundations.

Bents.

Bearings.

Expansion joints and truss intermediate hinge.

The capacity curves of the foundation and bents are compared

with the demands from inelastic Response History analysis (RHA).Similarly, the capacities of all bridge bearings, including pinned


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a

b

c

Fig. 13. OpenSees abutment model and soil profile: (a) three-dimensional simulation abutment model; (b) pushover analysis results in the longitudinal direction; (c)

abutment soil profile.

and sliding bearings, are compared with the maximum demandsfrom inelastic RHA. Finally, the opening and closing of the jointgaps is monitored to allow for predicting the actual behaviorof the structure and a possible bounding between differentstructural segments and any slippage at expansion bearings. Anindication of structural damage is considered if the superstructuredisplacement in the longitudinal direction frequently exceeded thegap provided at the expansion joints, causing pounding betweenbridge segments. Total collapse is assumed if the displacementexceeded the seat width.

As mentioned above, two analysis platforms are employedin the present study. Additionally, two analytical models weredeveloped using SAP2000; a detailed and a simplified model.

The latter idealization simplifies modeling of the superstructure,while refines the substructural members idealization, whichare responsible for dissipating energy. This allows reducing thesize of the simplified model, which is transferred to ZEUS-NL for inelastic analysis. Bearings friction and expansion jointsare considered only in the ZEUS-NL refined fiber model, whichaccounts for the material inelasticity and geometric nonlinearity.The kinematic soilstructure interaction is accounted for in allmodeling approaches by restraining the pile caps and the caissonsat their center of gravity with grounded springs representing thefoundation stiffnesses estimated from the extensive OpenSeesinelastic analyses. Three different analytical models are thereforeemployed in the present study. The following analyses are carriedout using the three structural models:

Free vibration analyses using the detailed and a simplifiedSAP2000 models as well as the ZEUS-NL model.

Inelastic static pushover analyses in the longitudinal andtransverse directions using the ZEUS-NL model.

Elastic dynamic analyses using the SAP2000 and the ZEUS-NLmodels.

Inelastic dynamic analyses in the longitudinal and transversedirections using the ZEUS-NL models and 144 site-specificsynthetic records representing three seismic scenarios.

Eigenvalue analyses are first conducted to determine the non-cracked horizontal and vertical periods of vibration and modeshapes. This simple analysis is used as an initial validation toolof the analytical models. Inelastic static pushover analyses areperformed in both the longitudinal and transverse directions

of the bents and the bridge foundations to evaluate theirlateral capacities. Elastic and inelastic timehistory analyses areinitially performed to verify the analytical models and tune thenon-hysteretic damping level. Extensive inelastic RHA are thenexecuted toexamine theresponseof thestructures under theeffectof site-specific input ground motions with increasing severity.Sample results of these comprehensive analyses are discussedbelow.

6. Dynamic characteristics and verifications of the analytical

models

Sample results from the SAP2000 and ZEUS-NL free vibrationanalyses are presented in Figs. 14 and 15. The mode shapes and

the corresponding periods of vibration obtained from the SAP2000and ZEUS-NL models are presented. It is important to note that the


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(a) Detailed model. (b) Simplified model.

Fig. 14. Comparison of the bridge dynamic characteristics obtained from the detailed versus the simplified SAP2000 models.

(a) ZEUS-NL model (without reinforcement). (b) SAP2000 simplified model.

Fig. 15. Comparison of the bridge dynamic characteristics obtained from the ZEUS-NL and SAP2000 models.

bridge bearings frictional resistance is disregard in these analyses.The significance of bearing friction on the dynamic characteristicsof highway bridges was confirmed from several previous studies(e.g. [28]). Multi-linear elasto-plastic joint elements are thereforeemployed to idealize the movable bearings in the refined inelasticanalyses carried out using ZEUS-NL, as explained above. The freevibration analyses confirmed that about 350 modes of vibrationare required to reach a 90% mass participation in the two principle

directions of the bridge. Higher modes of vibrations notablycontribute to seismic response due to the length of the 2164-

meter (7100-foot) bridge and the non-uniform distribution ofstiffness and mass of this complex bridge. High longitudinalvibrations are observed in the truss portal frames from the simplefree vibration analyses, which reflect the vulnerability of thesestructural members.

To verify the simplifications made in the superstructure model-ing, the detailed and the simplified SAP2000 models are comparedin Fig. 14. It is clear that the first mode shapes and their periods of

vibrationare comparable from thetwo models, which lends weightto the simplifications adopted in the superstructure idealization.


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Table 4

Surface input ground motions used for assessment of the Caruthersville Bridge.

Scenario Ground motion seta Return period (year) Site effectsb Asynchronous ground motionsc Number of records

1

Rec1-500-SE

500 12

Rec2-500-SE 12

Rec3-500-SE 12

2

Rec1-500-AS

500

12

Rec2-500-AS 12Rec3-500-AS 12

3

Rec1-1000-AS

1000 12

Rec2-1000-AS 12

Rec3-1000-AS 12

4

Rec1-2500-AS

2500 12

Rec2-2500-AS 12

Rec3-2500-AS 12

a Each input ground motion set has twelve different ground motion records, corresponding to the six soil profiles at Bents 1, 11, 15, 19, 22 and 27 in the longitudinal and

transverse directions of the bridge.b Only site effects were accounted for.c Incoherency, wave passage and site effects were accounted for.

Moreover, the SAP2000 simplified model is compared with ZEUS-

NL fiber model in Fig. 15. The effect of steel reinforcement, whichincreases stiffness and shortens periods of vibrations, was omittedfrom the latter model to obtain rational results from the two mod-els. Theresults reflect theeffectiveness of transferring theSAP2000model to ZEUS-NL and confirm the rationality of the latter modelfor predicting the inelastic response of the bridge. Adding the steelreinforcement to the model increases stiffness and shortening theperiods of the bridge by up to 8%.

7. Comparison of seismic demand versus capacity

Inelastic static pushover analyses are performed in both thelongitudinal and transverse directions of bridge to evaluate thelateral capacity of the bridge bents. This analysis is also conducted

for all foundation classes to estimate their capacities, as explainedabove. Extensive inelastic RHAs are finally executed using site-specific input ground motions to examine the response of thebridge under various seismic scenarios with increasing severity.Capacities of the foundations, bents, bearings and expansion jointsare compared with seismic demands at various hazard levels.Sample results from these comprehensive analyses are discussedbelow to highlight the significance of advanced simulationapproaches in identifying areas of vulnerability of complex bridges.Detailed results from the comprehensive seismic assessment studycarried out for the A-1700 Bridge under the effect of differentseismic scenarios are presented elsewhere [25,29,30].

7.1. Input ground motions

Probabilistic Seismic Hazard Analyses (PSHA) for hard rock siteconditions were performed for the bridge site [31]. Three hazardlevels corresponding to return periods of 500, 1000 and 2500 yearswere considered. Fig. 16(a)shows thelocation of theCaruthersvilleBridge with respect to the New Madrid seismic faults, while thedeveloped Uniform Hazard Spectra (UHS) are shown in Fig. 16(b).Ten records of spectrum compatible ground motions for site classA were generated for each of the three hazard levels of 10%,5% and 2% probability of exceedance in 50 years. Each of therecords included two components aligned in the longitudinal andtransverse directions with respect to the Caruthersville Bridge.For more details about seismic hazard analysis of the bridge site,reference is made to Ferenndez and Rix [31]. Three input ground

motions were selected from the above-mentioned ten recordsfor propagation through the thick embayment deposits [32]. Six

distinct soil profiles along the length of the bridge were also

selected for use in one-dimensional seismic site response analyses.Therefore, each of the three input ground motions used in thesite response analyses produced twelve different ground motionrecords, corresponding to the soil profiles at Bents 1, 11, 15, 19,22 and 27 (refer to Fig. 5) in the longitudinal and transversedirections of the bridge. The soil profile, shear wave velocity,and modulus reduction and damping curves that were used forsite amplification are presented elsewhere [30]. Fig. 16(c) depictsthe response spectrum of an input ground motions generatedfor bedrock in the longitudinal direction of the bridge and thosepropagated to the surface for the 2500 years seismic scenario. It isclear that thesurface motions have lower amplification in theshortperiod range compared with bedrock motions. The former groundmotions significantly amplify in the mid period range, which is inthe range of 0.11.0 s for the 500 years records and above 1.0 s forthe 1000 and 2500 years return period records [32].

Given the length of the 2164-meter (7100-foot) bridge, theground motion spatial variability were also included in thepropagated records. In addition to local site effects, two additionaleffects were considered [33,3,34]: (i) geometric incoherenceeffects due to the scattering in the heterogeneous ground, and (ii)wave-passage effects, where non-vertical waves reach differentpoints on the ground surface at different times, producing a timeshift between the motions at those points. The coherency functionby Abrahamson et al. [33], which provides quantitative measureof similarities in two ground motions recorded at differentseismic stations, was applied to compute incoherency at Bents1, 11, 15, 19, 22 and 27 [34]. The input ground motions werepropagated to the ground surface through the corresponding six

soil profiles after including the geometric incoherence and wave-passage effects. Four different scenarios of ground motions weregenerated following the above-mentioned procedures. Table 4shows the surface input ground motions used for assessment ofthe Caruthersville Bridge. In total 144 input ground motions areused in the present assessment study. Ground motions due to siteeffect only are used since they represent the typical site-specificground motions used in inelastic dynamic analysis. Spatiallyvariedmotions due to site effect, wave passage effect and ground motionincoherency are also employed since they represent the realearthquake effect. The development of spatially variable groundmotions is discussed in more detail by Tsai and Hashash [34] andMwafy et al.(2010). The latterstudyalso investigatedthe impactofwave passage and ground motion incoherency, and concludedthatasynchronous ground motions have significant implications on theseismic behavior of extended bridges.


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a

b c

Fig. 16. Site-specific input ground motions: (a) location of the Caruthersville Bridge; (b) bedrock Uniform Hazard Spectra (UHS) for return periods of 500, 1000 and

2500 years; (c) input ground motions generated for bedrock and those propagated to the surface for the 2500 years seismic scenario.

(a) Capacity of expansion joints vs displacement demand (most critical

ground motion set in longitudinal direction).

(b) Capacity vs demand at the top columns of Bents 1921 (ground

motions are in transverse direction).

(c) Bearing force capacity versus demand (most critical ground motion set in transverse direction).

Fig. 17. Capacity versus demand of different bridge components under the 500 years earthquake scenario.

7.2. Capacity versusdemandunder the 500yearsearthquake scenario

Due to the length of the bridge and the non-uniform distribu-

tion of stiffness and mass, higher modes of vibrations notably con-tribute to the seismic response. Inelastic response history anal-

yses carried out in the transverse directions of the bridge indi-cate that the drift demands are acceptable (less than 0.63%), withthe exception of the high displacement at the truss intermediate

hinge. Ground motion set 2 (Rec2-500-AS) significantly amplifiesthe truss deformation demands compared with other records. The


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(a) Mapping of plastic hinge formation (most critical ground motion set in transverse direction).

(b) Capacity versus bending moment demands of foundation class 3(ground motions are in transverse direction).

(c) Bearing force capacity versus demand (most critical ground motion set in longitudinal direction).

Fig. 18. Response of different bridge components under the 1000 years earthquake scenario.

relative displacements between piersalong the length of the bridgeare inconsistent as a result of the difference in stiffness, whichcauses high demands on the superstructure. Higher drift demandsareobserved in the longitudinal direction of the bridge (1.1%) com-pared with the transverse direction due to its lower stiffness. Highdeformations are observed in the truss portal frames, which arealso observed from free vibration analysis. The high relative dis-placement demands in the longitudinal direction cause poundingat the two abutments and the expansion joint of Bent 8, as shownin Fig. 17(a). The response of the foundation system are accept-able in both directions, while high demands are observed at thetop columns of Bents 1521, as shown in Fig. 17(b). The demandsof other bents are moderate to high. Several plastic hinges areobserved when applying the load in the transverse direction, par-

ticularly at the top columns of Bents 1521. The non-uniform dis-tribution of stiffness along the height of these piers causes high

stress concentrations at the top columns. A number of bridgebearings are vulnerable under the effect of this earthquake sce-nario, particularly the bearings at the expansion joints, as shownin Fig. 17(c).

7.3. Capacity versus demand under the 1000 years earthquake

scenario

The displacement and force demands increase under thisseismic scenario by up to 200% and 100%, respectively, comparedwith the 500 years seismic scenario. The maximum drift demandsin the longitudinal (3.8%) and the transverse (2%) directions areunacceptable. Exceeding the 3% drift limit is considered in the

present study as an indication of extensive structural damage. Thehigh relative displacement demands in the longitudinal direction


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Fig. 19. Bearing capacity versus demand under the 2500 years ground motions (most critical record in the longitudinal direction).

cause pounding at all expansion joints. Unacceptable displacementdemands (934 mm) are observed in the transverse direction at thetruss intermediate hinge, particularly under the effect of Groundmotion set 2 (Rec2-1000-AS). Extensive damage and yielding arealso detected in several bents. The high ductilitydemands imposedon the top columns of Bents 1521 cause severe damage. Thespread of plastic hinges is extensive, as shown in Fig. 18(a), while asevere damage in several bearing is confirmed, as confirmed from

Fig. 18(c). Unacceptable inelastic response is observed in somefootings under the effect of the 1000 years seismic scenario, asshown in Fig. 18(b).

7.4. Capacity versus demand under the 2500 years earthquake

scenario

Displacement demands significantly increase in the longitudi-nal direction of the bridge by up to 90% compared with those fromthe 1000 years input ground motions. The results obtained fromGround motion set 2 (Rec2-2500-AS) confirm that the drift at thetop of bents (7.25%) considerably exceeds the collapse limit state.The very high relative displacement demands cause collision be-tween all bridge segments. Significantly higher force demands are

also observed in the bents and the foundation system comparedwith those observed from the 1000 years earthquake scenario. Theinelasticresponse of the foundation systemis confirmed, while ex-tensive damage and yielding at the base of major bents supportingthe steel truss are detected. Most of the bridge bearings are vul-nerable (capacity/demand ratio is less than unity) under the effectof the 2500 years earthquake scenario, as shown in Fig. 19. The re-sults confirm the need to retrofit different bridge components tomitigate the potential seismic risk.

8. Conclusions

This paper presents highlights of a project carried out toassess the seismic response of a 59-span bridge, considering

SoilStructure Interaction (SSI). The focus is on describing themethodology adopted to idealize thebridge andits foundation sys-tem, while only sample results from the extensive elastic and in-elastic analyses under the effect of 144 site-specific input groundmotions are presented. The bridge was built in the vicinity of amajor source of earthquakes; the New Madrid Seismic Zone, andincludes typical deficiencies of bridges constructed without cur-rent seismic provisions. The refined three-dimensional simulationsof the bridge and its foundation system carried out using veri-fied analysis tools and a detailed site-specific seismic hazard studyenabled the identification of areas of vulnerability of the investi-gated bridge and assessment of its response at three hazard lev-els, corresponding to 500, 1000 and 2500 years return periods. Thestudy confirmed that simplifying modeling assumptions of differ-

ent bridge components may have significant impact on the seismicresponse of complex bridges. Higher modes of vibration notably

contributed to the seismic response of the I-155 bridge due to thelength of the bridge and the non-uniform distribution of stiffnessandmass.Under the500 years groundmotions,the response of thebridge was unacceptable due to the observed yielding and damagein a number of bents and bearings. The demands correspondingto the 1000 years ground motions almost exceeded the collapselimit state and the capacity of bridge components. Indications ofyielding in foundations were also observed. The displacement and

force demands under the effect of the 2500 years earthquake sce-nario significantly increased compared with the 1000 years inputground motions and exceeded by far the collapse limit states. Thepresented assessment study confirmed the need to retrofit differ-ent bridge components to mitigate potential seismic risk. The toolsand procedures used for this assessment are applicable to similarsituations of complex bridges constructed in a mix of material andstructural systems.

Acknowledgements

This study was funded by the Missouri Department ofTransportation (MoDOT) with Jacobs Civil Inc. as main contractor,and the Mid-America Earthquake Center (MAE) as sub-contractor.

The MAE Center is an Engineering Research Center funded bythe National Science Foundation under cooperative agreementreference EEC 97-01785. All findings and conclusions are those ofthe authors and do not represent Jacobs or MoDoT.

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