Seismic assessment of an existing non-seismically designed major bridge-abutment–foundation system PDF

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Engineering Structures 32 (2010) 2192–2209 Contents lists available at ScienceDirect Engineering Structures journal homepage: www.elsevier.com/locate/engstruct Seismic assessment of an existing non-seismically designed major bridge-abutment–foundation system Aman Mwafy a , Oh-Sung Kwon b,∗ , Amr Eln...

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Engineering Structures 32 (2010) 2192–2209

Contents lists available at ScienceDirect

Engineering Structures journal homepage: www.elsevier.com/locate/engstruct

Seismic assessment of an existing non-seismically designed major bridge-abutment–foundation 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 Emirates

b

Department of Civil, Architectural and Environmental Engineering, Missouri University of Science and Technology, Rolla, MO 65409, USA

c

Department of Civil and Environmental Engineering, University of Illinois at Urbana-Champaign, Urbana, IL 61801, USA

article

info

Article history: Received 14 May 2009 Received in revised form 13 March 2010 Accepted 15 March 2010 Available online 18 April 2010 Keywords: Major highway bridges Seismic assessment Inelastic dynamic analysis Soil–structure interaction High seismicity regions Bridge deficiencies

abstract A comprehensive study carried out to assess the seismic response of a 59-span bridge using a refined inelastic modeling approach and considering Soil–Structure Interaction (SSI) is summarized in this paper. The focus is on describing the methodology adopted to idealize the bridge and its foundation system, while only highlights from the extensive elastic and inelastic analyses are presented. The bridge represents a typical case of vulnerable complex bridges since it was built in the early seventies with minimal seismic design requirements at a distance of about 5 km from a major fault. The SSI analysis is significant in this study due to the length of the bridge, the massive and stiff foundation, and the relatively soft deep soil of the site. A series of three-dimensional dynamic response simulations of the entire bridge are conducted using several analysis tools to verify the developed analytical models. The performance-based assessment 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 the 2164-meter bridge at various hazard levels. It is concluded that the seismic response of the bridge at the 500 years ground motions does not meet today’s standards, while the demands under the effect of the 1000 years ground motions almost exceed the capacity of most bridge components. The demands significantly increase under the effect of the 2500 years earthquake scenario and considerably exceed the collapse limit states. The results clearly reflect the benefit of retrofitting different bridge components to mitigate the anticipated seismic risk. The presented assessment study contributes to improve public safety by exploiting the most recent research outcomes in predicting the seismic response of complex highway 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 to significant changes in seismic design provisions of bridges. The introduction 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 uniform safety for different types of bridge systems. The release of the FHWA Retrofitting Manual for Highway Structures [3] also provides comprehensive procedures for assessment and retrofitting highway bridges based on recent experiences in the US, Japan, and other countries. This major revision in the FHWA Seismic Retrofitting Manual introduces a performance-based retrofit philosophy and defines several methods for the detailed evaluation 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).

0141-0296/$ – see front matter © 2010 Elsevier Ltd. All rights reserved. doi:10.1016/j.engstruct.2010.03.022

and retrofitting guidelines draw attention to the need for seismic assessment of complex highway bridges designed to preceding provisions to determine the level of risk associated with loss of serviceability or possible damage. This is particularly significant in the light of the continuous updates in seismic hazard maps for several regions (e.g. [4]). A large number of bridges were designed and constructed at a time when bridge codes were insufficient according to current standards. The deficiencies in highway bridges designed prior to the seventies result in excessive seismic displacements and large force demands that were substantially underestimated. The anticipated damage includes unseating and pounding of superstructure at abutments and expansion joints, shear and flexural failure of RC piers, beam–column joint failure, footing and abutment failure and amplification of response due to soft/liquefiable soil (e.g. [5]). The existing bridge inventory designed to previous provisions is thus likely to suffer damage when subjected to seismic scenarios comparable to those observed in severe earthquakes (e.g. San Fernando, USA, 1971; Loma Prieta, USA, 1989; Northridge, USA, 1994).

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In-depth performance-based seismic assessment studies of major highway bridges point the way towards improving the understanding of the complex seismic performance of similar structures. This performance-based evaluation approach requires bridges to satisfy different performance criteria for different levels of ground motion. For instance, the bridge may suffer minor damage but should be operational under frequent earthquakes with low intensity. Under infrequent earthquakes with large intensity, the bridge should provide an acceptable level of life-safety. Quantifying the level of risk associated with anticipated earthquake scenarios enables taking rational decisions to retrofit, replace or accept the risk. The preliminary screening and prioritization approach proposed by the FHWA Retrofitting Manual for Highway Structures [3] classifies bridges according to their Seismic Retrofit Categories (SRC) to identify minimum requirements for screening, evaluation and retrofitting. The FHWA Retrofitting Manual recommends performance criteria according to bridge importance and anticipated service life, with more rigorous performance required for important, relatively new bridges, than for standard bridges at the end of their service life. The manual also introduces a performance-based retrofit philosophy similar to that used for the performance-based design of new buildings and bridges. Following the preliminary screening process, the FHWA retrofit manual [3] recommends six evaluation procedures, with increasing order of rigor, for quantitatively assessing the seismic performance of existing highway bridges. Method E, which is based on the inelastic time history analysis, compares seismic demands with member capacities and determines the degree of damage to the structure at the desired earthquake levels. The latter method, which requires considerable computational effort and a significant level of skill to interpret the results, is recommended for irregular complex bridges with the potential for substantial inelastic behavior particularly when sitespecific ground motions are used. The major bridge assessed in the present study requires the most rigorous evaluation procedure recommended by the FHWA retrofitting manual. The primary objectives of this paper are therefore as follows:

• Discuss the comprehensive methodology adopted to realistically assess the seismic response of a major bridge and its foundation system taking into account the most recent advances in the analytical modeling approaches of substructure, superstructure, abutments and foundations. • Present sample results from the comprehensive vulnerability study to highlight the significance of advanced simulation approaches in identifying areas of vulnerability of complex bridges. • Investigate the significance of the recent understanding about the seismicity of the site of the investigated bridge in providing insight 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-155 across the Mississippi River between Pemiscot County, Missouri, and Dyer County, Tennessee, is a major bridge that has a pressing need for detailed vulnerability assessment [6]. Although the 59-span 2164-meter bridge is about 5 km from a presumed major fault, it was constructed in the early seventies with minimal seismic design requirements by today’s standards. According to the design information [7], Peak Ground Accelerations (PGA) of 0.10g and 0.06g were employed in the seismic design of the bridge in the transverse and longitudinal directions, respectively. Clearly this reflects the limited knowledge about the seismicity of the construction site at the time of designing the bridge; the whole field of study has advanced since the time of construction. The superstructure consists of eleven units separated by expansion joints and supported on a variety of elastomeric and steel bearings. The

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main channel crossing is composed of two-span asymmetrical cantilever steel truss and ten-span steel girders, while approach spans are precast prestressed concrete girders. The substructure includes piers on deep caissons and bents on steel friction piles driven into the near surface silty sands and clayey materials. Bedrock is located more than 800 m below the sand, gravel, and hard clay strata. Fig. 1 shows a three-dimensional view of the I-155 bridge and pictorially summarizes the adopted three-dimensional analytical modeling approach, which is discussed in the following sections. This brief description of the Caruthersville Bridge highlights the pressing need to reliably assess its response under anticipated seismic scenarios. 3. Development of simulation models for inelastic analysis Detailed three-dimensional dynamic response simulations of the entire bridge including foundations and soil effects are carried out using a number of verified analysis platforms. The finite 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. The SAP2000 analytical models are mainly employed for verifications of the ZEUS-NL fiber model before executing the extensive inelastic analysis. ZEUS-NL is mainly employed to estimate the capacities and demands from inelastic pushover and response history analyses. The latter finite element analysis platform was developed at Imperial College London and at University Illinois at Urbana–Champaign, and has been extensively employed in large projects (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 represent all sub- and super-structural components for elastic analysis. Steel and concrete cross-sections from the SAP2000 library are employed to realistically model different structural members for elastic analysis. This modeling approach, particularly for the superstructure, is computationally demanding for inelastic response history analysis. Also, the design philosophy of bridges relies on bridge piers to dissipate energy rather than the superstructure, which remains elastic. The detailed SAP2000 analytical model is therefore modified to reduce the number of elements and nodes to a manageable limit for inelastic analysis. With the exception of spans with truss, the superstructure is replaced with a number of cross sections with equivalent geometrical properties connected together 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 subdividing the columns to a number of elements to accurately monitor the inelastic response during time–history analysis. Moreover, the SAP2000 joint constraints, which are not available in ZEUS-NL, are replaced with strong arms. The simplified SAP2000 model was transferred to ZEUS-NL for inelastic analysis. Due to the complex behavior of the truss, it was transferred to ZEUS-NL without simplifications. Based on available cross-sections in the library of SAP2000 and ZEUS-NL, equivalent cross-sections are adopted for modeling of the bridge members. In the detailed ZEUS-NL model, each structural member is assembled using a number of cubic elasto-plastic elements capable of representing the spread of inelasticity within the member cross-section and along the member length via the fiber analysis approach. Sections are discretized to steel, confined

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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 stress–strain response at each fiber is monitored during the entire multi-step analysis. Therefore, the degradation of stiffness with increasing inelasticity is fully accounted for in the ZEUS-NL inelastic analysis. Equivalent gravity loads and mass are distributed in the ZEUSNL fiber model on the superstructure and along the piers height. The employed distributed mass elements utilize cubic shape function and account for both the translational and rotational inertia. The total weight of the bridge is 1562,549 kN (351,275 kip), which includes superstructure, substructure, non-structural members, pile caps and caissons. 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 the steel girders and the truss spans. This is due to the lower weight of the steel members compared with the concrete counterparts and the lightweight deck of the steel and truss spans. Also, the massive weight of the caissons significantly increases the substructure weight 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 bridge and the lack of reliable information confirming the actual material characteristics, nominal material properties are used in analysis. The cylinder concrete strength (fc0 ) is 21 MPa (3000 psi) for normal concrete and 28 MPa (4000 psi) for prestressed concrete. The

yield strength of reinforcing steel (fy ) is 40,000 psi. Structural carbon steel ASTM A36-69 (fy = 248 MPa or 36,000 psi) is used for truss members and steel beams. A bilinear model was used to idealize steel members and reinforcement. In this model, the loading and unloading in the elastic range follow a linear function throughout various loading stages with constant stiffness represented by the Young’s modulus of steel. In the post-elastic range, a kinematic hardening rule for the yield surface defined by a linear relationship is assumed [13]. A uniaxial constant ‘active’ confinement concrete model was employed in the ZEUS-NL analytical model. This model has a good balance between simplicity and accuracy and includes enhancements in the cyclic degradation rules, inelastic strain, and shape of unloading branches [14]. The model, which incorporates the influence of confinement effects on the peak stress and strain as well as on the post-peak stress–strain relationship, can provide a good estimation of the cyclic response of RC members under cyclic and dynamic loading. As suggested from the bridge drawings, appropriate translational (∆) and rotational (r) Degrees of Freedom (DOFs) are released at the wind transfer device of the truss, as shown from Figs. 1 and 2. The translational DOFs, ∆X , and the rotational DOFs, rY and rZ , of the stringers and the deck are also released at certain expansion joints as indicated in the bridge drawings. Bridge bearings and expansion joints are realistically modeled using ZEUS-NL joint elements. The rotation about the transverse axis, rY , is only

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(a) Released DOFs at the truss intermediate hinge.

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(b) Gap elements controlling displacement.

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

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(a) ZEUS-NL model of the expansion shoe at Bents 15, 19 (left), 21 (right) and 25.

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(b) ZEUS-NL model of the Bronze self-lubricating bearings at Bents 8, 14 (left), 26 (right), 31, 41 and 51.

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

allowed at the fixed shoes. Based on the drawings, these bearings are located at the substructure–superstructure connection of Bents 2–7, 9–13, 14 (right), 16–18, 20, 21 (left), 22–24, 26 (left), 27–30, 32–40, 42–50 and 52–59. On the other hand, the sliding bearings are located at the two abutments and at the substructure–superstructure 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 Mander et al. [15]. This bearing type is highly susceptible to instability since the horizontal motion in the longitudinal direction is accommodated through a rocking motion. This motion causes friction resulting from a rolling resistance at the base and a Coulomb friction at the top hinge [16,17]. Mander et al. [15] observed that debris and uneven wear cause the experimental hysteresis loops to be irregular. The ZEUS-NL model of this bearing is shown in Fig. 3(a). The initial stiffness of this bearing, K1 , is 14 kN/mm with a post-yield stiffness, K2 , of 0.018k1 . The frictional coefficient is 0.04. The analytical model of the expansion bearing at Pier 19 (right) has also an initial stiffness, K1 , of 14 kN/mm with a perfectly plastic post-yield stiffness. However, the longitudinal displacement of this bearing is restricted to 6.0 in. The frictional coefficient is 0.20, as recommended for the low-type expansion bearings investigated by Mander et al. [15]. Movable bearings, which typically have small friction coefficient at the low velocity rates, may have higher friction under high

seismic deformation. For un-lubricated elastomeric bearings, this coefficient at high velocities ranges from 5% to 15%, or even higher at the low temperature (e.g. [18]). It was also concluded in previous experimental studies that the coefficient of friction slightly decreases 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). The initial stiffness, K1 , is taken equal to 123 kN/mm based on the low type sliding bearings tested by Mander et al. [15]. The frictional coefficient of the Bronze Self-lubricating surface is taken equal to 0.10. The post yield stiffness is estimated from the geometry of the system. The horizontal movement of this bearing implies a vertical uplift of the superstructure. The post yield stiffness, K2 = W /r, where W is the weight and r is the radius of the bearing surface curvature. The top plates of these bearings are slotted to allow for 3.5 in. of movement. The displacement in the longitudinal direct...