Performance-Based Design of Seismically Isolated Bridges A

Performance-Based Design of Seismically Isolated Bridges

by

Eric Leonard Anderson

B.S. (California Polytechnic State University) 1987M.S. (University of California, Berkeley) 1996

A dissertation submitted in partial satisfaction of the

requirements for the degree of

Doctor of Philosophyin

Engineering - Civil and Environmental

in the

GRADUATE DIVISION

of the

UNIVERSITY OF CALIFORNIA, BERKELEY

Committee in charge:

Professor Stephen A. Mahin, ChairProfessor Gregory L. Fenves

Professor Ole Hald

Spring 2003

The dissertation of Eric Leonard Anderson is approved:

University of California, Berkeley

Spring 2003

Chair Date

Date

Date


Copyright 2003

by



by


B.S. (California Polytechnic State University) 1987M.S. (University of California, Berkeley) 1996

A dissertation submitted in partial satisfaction of the

requirements for the degree of

Doctor of Philosophyin

Engineering - Civil and Environmental

in the

GRADUATE DIVISION

of the

UNIVERSITY OF CALIFORNIA, BERKELEY

Committee in charge:


Professor Ole Hald

Spring 2003

The dissertation of Eric Leonard Anderson is approved:


Spring 2003

Chair Date

Date

Date


Copyright 2003

by


i

Table of Contents

Section Page

Chapter 1 Introduction 1

1.1 Seismic Protective Systems for Bridges 1

1.2 Commentary on the State of the Art 4

1.3 The Protective Systems Research Program 6

1.3.1 Overall Objectives of the Program 6

1.3.2 Overall Tasks Undertaken 6

1.3.3 Protective system devices considered 7

1.3.4 Test Models 8

1.4 Scope of this Report 9

1.5 Organization of this Report 10

Chapter 2 Preliminary Evaluations 16

2.1 Introduction 16

2.2 AASHTO Guide Specifications 16

2.3 Parametric Seismic Analysis of Idealized Elevated Isolated Bridges19

2.3.1 Introduction 19

2.3.2 Idealized Modeling Assumptions 20

2.3.3 Analytical Models 22

2.3.4 Ground Motion Time-Histories 26

2.3.5 Parametric Studies 27

2.3.5.1 Influence of Substructure or Initial System Flexibility 27

2.3.5.2 Influence of Isolator or Second-Slope Flexibility and Characteristic Strength 31

ii

2.3.5.3 Influence of Substructure Mass 36

2.3.5.4 Influence of Substructure Damping 39

2.3.5.5 Influence of Bi-Directional Ground Motion Input 43

2.3.5.6 Influence of Superstructure Mass Eccentricity 45

2.3.5.7 Influence of Substructure Stiffness Eccentricity 48

2.4 Summary 51

Chapter 3 Testing of a Bridge Deck Model 67

3.1 Introduction 67

3.1.1 Similitude Requirements 68

3.1.2 Design criteria 70

3.1.3 Design Development 72

3.1.4 Final Specimen Design 73

3.1.5 Test Set-Up and Protocol 77

3.2 Earthquake Histories for Testing 85

3.3 Pre-Test Analysis of Bridge Deck Model 102

3.4 Experimental Results for Configuration 1 104










3.14 Experimental Characterization Data for FP Slider Composites 122

iii

Chapter 4 Evaluation of Bridge Model Test Data 127

4.1 Introduction 127

4.2 Influence of Substructure Flexibility 128

4.3 Influence of Isolator Second-Slope Flexibilityand Strength 130

4.4 Influence of Substructure Mass 133

4.5 Influence of Bi-directional Motions 135

4.6 Influence of Substructure Strength 136

4.7 Influence of Superstructure Mass Eccentricity 138

4.8 Influence of Substructure Stiffness Eccentricity 140

4.9 Response of Two-span Isolated Bridge 142


4.9.2 Characteristic Distribution of Force & Displacement Demands 143

4.9.3 Local Geometric Effects 144

4.10 Substitute System vs. MDOF Response 146

4.11 Influence of Ground Motion 149

4.12 Influence of Vertical Motions on Sliding Systems 154

4.13 Effect of Impact Against Bearing Restraint System 155

4.14 Effect of Bearing Wear on Systems EmployingFP Bearings 157

Chapter 5 Displacement Estimates in Isolated Bridges 214


5.1.1 Problem Summary 214

5.2 Analytical Bridge Models 215

5.3 Ground Motion Time-history Suites 216

iv

5.3.1 Suites for Evaluation of the AASHTO Uniform Load Methodand R-factor Provisions 218

5.3.2 Suites for Evaluation of the Effect of Bi-directional Motionson AASHTO Provisions 226

5.3.3 Suites for Parametric Studies of Isolated Bridge Responseand Influence of Substructure Mass and Damping 239

5.4 An Evaluation of the AASHTO Uniform Load Method 250


5.4.2 AASHTO Uniform Load Method 250

5.4.3 Evaluation Procedure 254

5.4.3.1 Method A: Design Spectrum 254

5.4.3.2 Method B: Specified Ground Motion 255

5.4.4 Isolated Bridge Systems 257

5.4.5 Numerical Analysis Procedure 257

5.4.6 Results 258

5.4.7 Observation summary 269

5.5 Bi-Directional Earthquake Shaking 273


5.5.2 Isolated Bridge System Models 277

5.5.3 Analytical Procedure and Response Parameters 278

5.5.4 Results 280

5.5.5 Observation Summary 283

5.6 Influence of Substructure Yielding on System Response 288


5.6.2 System Characterization 290

5.6.3 AASHTO Substructure Component Design Procedures 290

5.6.4 Evaluation Procedure 293

v

5.6.4.1 Response parameters 295

5.6.5 Results 296

5.6.5.1 Displacement and Base Shear Demand 296

5.6.5.2 Substructure Ductility Demand 298

5.6.5.3 Ratio of Peak Response for Systems with and without YieldingSubstructures 299

5.6.5.4 Displacement Distribution Ratios for Systems with and withoutYielding Substructures 301

5.6.5.5 Time-history variations with and without Yielding Substructures:Displacement Redistribution 304


5.7 Parametric Study of Nonlinear Isolated Bridge Response 308


5.7.2 System Characterization 310


5.7.4 Analytical Procedure 311

5.7.5 Results 312

5.7.5.1 Displacement and Base Shear Response 312

5.7.5.2 Isolator and Substructure Displacement Response 319

5.7.5.3 Isolator & Substructure Displacement Distribution Ratios 329

5.7.6 Conclusions 333

5.7.6.1 Total Displacement 334

5.7.6.2 Base Shear 336

5.7.6.3 Isolator Displacements 339

5.7.6.4 Substructure Displacement 340

5.7.6.5 Isolator/Substructure Displacement Ratios 341

5.7.6.6 Earthquake Magnitude and Distance 343

vi

5.7.6.7 Summary 344

5.8 Influence of Substructure Mass and Damping 346



5.8.3 Analytical Procedure and Response Parameters 349

5.8.4 Results: Effect of Substructure Damping 351

5.8.4.1 Total and Isolator Displacement Response 351

5.8.4.2 Substructure Displacement and Base Shear Response 354

5.8.5 Results: Effect of Substructure Mass 358

5.8.5.1 Total Displacement Response 359

5.8.5.2 Isolator Displacement Response 362

5.8.5.3 Substructure Displacement and Base Shear Response 365


5.8.6.1 Effect of Substructure Damping 369

5.8.6.2 Effect of Substructure Mass 371

5.8.6.3 Concluding Remarks 372

Chapter 6 Conclusions and Recommendations 373


6.2 Earthquake Simulation Studies 374

6.2.1 Proof-of-concept 374

6.2.2 Sensitivity to characteristics of ground shaking 375

6.2.2.1 Ground motion type: far-field, near-fault, and soft-soil 375

6.2.2.2 Bi-directional motions 377

6.2.2.3 Vertical Motions 378

6.2.3 Effect of System Configuration 378

6.2.3.1 Substructure Flexibility 378

vii

6.2.3.2 Isolator Strength and Flexibility 380

6.2.3.3 Substructure Mass 380

6.2.3.4 Mass Eccentricity 381

6.2.3.5 Stiffness Eccentricity 382

6.2.3.6 Substructure Yielding 384

6.2.4 Characteristic Distribution of Force and Displacement Demands 384

6.2.5 Local Kinematic Effects 385

6.3 Implications for Design Practice 386

6.3.1 Basic AASHTO Design Equation 387

6.3.1.1 Reliability 387

6.3.1.2 Bi-directional Effects 392

6.3.1.3 Equivalent SDF vs. MDF behavior: Influence of Substructure Mass and Damping 396

6.3.1.4 Parametric Study of Nonlinear Isolated Bridge Response 401

6.3.1.5 Substructure Yielding 410

6.3.2 Alternative Methods 413

6.3.2.1 Equivalent Linear Procedures 414

6.3.2.2 Approximate Nonlinear Methods 415

6.3.2.3 Performance-Based Design 416

6.4 Future Research Need 420

6.5 Conclusion 421

Bibliography 422

Appendix A Bridge Deck Specimen Plans 426

Appendix B Pre-test Analysis 448

Appendix C Test Logs 463

Appendix D Test Data Summary Figures 495

viii

Appendix E AASHTO Uniform Load Method 681

Appendix F Bi-directional Motions 702

Appendix G Substructure Yielding 730

Appendix H Parametric Study 769

Appendix I Substructure Mass and Damping 881

1

Abstract


by


Doctor of Philosophy in Engineering - Civil and Environmental



Professor Ole Hald

The response of simple isolated bridge overcrossings to a variety of multi-dimensional

earthquake inputs was investigated by analysis and experiment. These studies were

undertaken as part of the Protective Systems Research Program at UC Berkeley sponsored

by the California Department of Transportation to establish an understanding of global

and local behavior characteristics, including the effect on response of variations in isolator

properties, pier flexibility, mass and strength, substructure damping, and global mass and

stiffness eccentricity. A versatile 1/4−scale bridge model on flexible piers, enabling one-

or two-span configurations, was tested on an earthquake simulator to validate the findings

of the analytical studies performed. Provisions of the current AASHTO Guide

Specifications for Seismic Isolation Design were also evaluated using the results of the

analytical and experimental investigations. Specifically, the application of the Uniform

Load Method to the design of a broad range of isolated bridge systems was studied. The

2

effect of mass and damping assumptions, bi-directional ground motion inputs, and

substructure strength on these procedures was also evaluated.

Several overriding observations were derived from these investigations. First, these

studies illustrated the considerable overall ability of seismic isolation to provide an

effective means of earthquake resistance in simple bridge overcrossings. The considerable

durability and robustness of these systems under seismic loading was illustrated through

multiple, varied simulations, including multi-dimensional inputs of far-field, near-fault,

and soft-soil ground motions. Secondly, analytical evaluations undertaken were able to

identify sensitivity of the nonlinear behavior of simple isolated bridge types to various

ground motion and structural characteristics. Lastly, these studies established several areas

in the current AASHTO Guide Specifications in need of further development.

• Foremost, it was shown that the Uniform Load Method does not provide uniform reli-ability, pointing to the need for the refinement of current Performance-Based Designprocedures.

• Further, possible improvements to the Guide Specification provisions to more accu-rately account for bi-directional effects were suggested.

• It was also shown that the present R-factor approach may not adequately control duc-tility demands in substructure components.

• Conditions were identified for which the effects of substructure mass and dampingmay require more accurate consideration than presently provided for by typicalapproximate linearized single-degree-of-freedom design procedures.

3

The abstract of Eric Leonard Anderson is approved:


Spring 2003

Chair Date

Date

Date

1

Abstract


by


Doctor of Philosophy in Engineering - Civil and Environmental



Professor Ole Hald

The response of simple isolated bridge overcrossings to a variety of multi-dimensional

earthquake inputs was investigated by analysis and experiment. These studies were

undertaken as part of the Protective Systems Research Program at UC Berkeley sponsored

by the California Department of Transportation to establish an understanding of global

and local behavior characteristics, including the effect on response of variations in isolator

properties, pier flexibility, mass and strength, substructure damping, and global mass and

stiffness eccentricity. A versatile 1/4−scale bridge model on flexible piers, enabling one-

or two-span configurations, was tested on an earthquake simulator to validate the findings

of the analytical studies performed. Provisions of the current AASHTO Guide

Specifications for Seismic Isolation Design were also evaluated using the results of the

analytical and experimental investigations. Specifically, the application of the Uniform

Load Method to the design of a broad range of isolated bridge systems was studied. The

2

effect of mass and damping assumptions, bi-directional ground motion inputs, and

substructure strength on these procedures was also evaluated.

Several overriding observations were derived from these investigations. First, these

studies illustrated the considerable overall ability of seismic isolation to provide an

effective means of earthquake resistance in simple bridge overcrossings. The considerable

durability and robustness of these systems under seismic loading was illustrated through

multiple, varied simulations, including multi-dimensional inputs of far-field, near-fault,

and soft-soil ground motions. Secondly, analytical evaluations undertaken were able to

identify sensitivity of the nonlinear behavior of simple isolated bridge types to various

ground motion and structural characteristics. Lastly, these studies established several areas

in the current AASHTO Guide Specifications in need of further development.

• Foremost, it was shown that the Uniform Load Method does not provide uniform reli-ability, pointing to the need for the refinement of current Performance-Based Designprocedures.

• Further, possible improvements to the Guide Specification provisions to more accu-rately account for bi-directional effects were suggested.

• It was also shown that the present R-factor approach may not adequately control duc-tility demands in substructure components.

• Conditions were identified for which the effects of substructure mass and dampingmay require more accurate consideration than presently provided for by typicalapproximate linearized single-degree-of-freedom design procedures.

3

The abstract of Eric Leonard Anderson is approved:


Spring 2003

Chair Date

Date

Date

1

1 Introduction

1.1 Seismic Protective Systems for Bridges

Various types of bearings are frequently used in bridges at the superstructure support

points. These bearings provide support for superstructure vertical loads, while allowing

thermal and long-term horizontal deformations to develop without transmitting significant

lateral forces into the substructure. The three most widely used types of bridge bearings

are pot, disc, and curved sliding bearings [Stanton et al., 1993]. These bearings are

generally designed to accommodate rotations about any axis. Horizontal translation is

accomodated by sliding along a standard stainless steel-PTFE interface. Examples of pot

and curved sliding bridge bearings are shown in Figure 1-1.

In addition to the above cited benefits of bearings in bridges, such bearings may provide

an ideal method to control forces and accommodate horizontal displacements imposed on

a bridge during an earthquake. During an earthquake, inertial loads acting on the

superstructure can induce large forces in the substructure and relative displacements

between the superstructure and the ground. The displacement variation between the

superstructure and foundation depends on: the characteristics of the ground motion; the

2

mass, stiffness, strength, and energy dissipation characteristics of the substructure; the

mass and flexibility of the superstructure; and the support conditions at the abutments and

foundations. For conventional bridges, piers and abutments are designed to accommodate

this displacement through a ductile mode of behavior. Properly reinforced concrete piers

can be designed to achieve this required displacement ductility. However, large ductile

deformations in these members spalls the concrete, yields the reinforcement in areas of

plastic hinge formation, and results in permanent lateral displacements. After a large

earthquake, damaged piers must be repaired or replaced, and the bridge and roadway may

have to be realigned. Higher mode contributions, support movements, and other

contributions to local stress concentrations may also cause unforeseen damage in

superstructure, substructure, or foundation components. This damage may be difficult to

assess and costly to repair.

As an alternative to conventional bridge design, seismic isolation bearings can be

employed. Such seismic protection devices are intended to control and accommodate

horizontal superstructure displacements without damage to bridge components (or the

bearings themselves) during an earthquake. Seismic protection devices provide engineers

with additional tools to achieve performance requirements for the safety and functionality

of bridges in the event of large earthquakes.

Seismic isolation bearings can accommodate large displacements while limiting force

transmission. Isolation bearings displacing on the order of 20 to 40 in. (500 to 1000 mm)

or more during an earthquake have been designed for bridges located close to major

3

earthquake faults. Representative examples of elastomeric and sliding seismic isolation

bearings are shown in Figure 1-2 and 1-3.

To prevent damage to these bearings and to limit damage to bridge hinges, expansion

joints, and abutments, it is important to control seismic displacement and force demands.

Thus, seismic isolation bearings typically exhibit highly nonlinear behavior, with

considerable hysteresis under cyclic loading. Bearings typically provide between 8 and 30

percent effective damping during seismic excitation, depending on the bearing design and

amount of bearing displacement. This nonlinear behavior, combined with hysteretic,

energy dissipation may effectively control peak superstructure displacements, and

substantially reduce demands on the substructure and superstructure compared to an

elastically designed structure.

Supplemental energy dissipation devices can also be added to seismically isolated and

conventional bridges to increase the total effective damping of the system and further

control response. Representative examples of a supplemental viscous damping device is

shown in Figure 1-4.

Seismic isolation bearings and supplemental dampers are examples of innovative and

practical technologies to enhance the performance of bridges during earthquakes. They

provide a practical method for protecting life and property in the event of a major

earthquake event.

4

1.2 Commentary on the State of the Art

Because of the potential benefits of seismic protective systems, considerable research

related to their use has been completed in the past ten years. An extensive summary of this

history is included in a related work by Whittaker et al. [Whittaker et al., 1998] conducted

as part of the Berkeley-Caltrans Seismic Protective Systems Program. As such, a review

of this literature is not provided herein.

Most studies to date have focused on the behavior of the isolation bearings or energy

dissipation devices themselves. Relatively few studies have addressed the overall

performance of seismically protected bridge systems. Even fewer studies have thoroughly

examined the effects of more than one component of earthquake ground motion on bridge

response.

Through a synthesis and analysis of these previous studies and the observation of

experienced bridge engineers, important issues impeding the application of this

technology to bridges have been identified. Foremost among these, perhaps, is the

potentially large displacement and force demands on seismically isolated bridges located

near major faults. Large amplitude, long duration velocity pulses characteristic of near-

fault ground motions have raised concerns regarding the displacement demands of

relatively weak systems undergoing large displacements. This behavior generally gives

the overall structure the appearance of having a long effective period of vibration. For this

reason, the applicability of seismic isolation for bridges situated on soft soils has also been

questioned. This is because site response effects for large ground motions tend to produce

significant displacements in the long period range. Codification of design guidelines for

5

the application of seismic isolation to bridge systems is also in its relative infancy. Few

seismically isolated bridges have been subjected to large design level earthquakes in the

field, or tested in the laboratorty, to confirm the adequacy of these guidelines. Many of the

concepts and procedures incorporated in these design methodologies have essentially been

adapted from the building code provisions [U.B.C., 1994], and as such, may not

adequately consider the differences in dynamic behavior and performance objectives

between isolated buildings and bridges.

The design of highway bridges with seismic isolation and supplemental damping is

specified in the American Association of State Highway and Transportation Officials

Guide Specifications for Seismic Isolation Design [AASHTO, 1999], referred to in this

report as the Guide Specifications or Guide Spec. The Guide Specifications provide

requirements for analysis of seismically protected bridges, isolation bearing and damper

design, and testing and acceptance of bearings and damper components. The Guide

Specifications incorporate a simplified linearized procedure for bridge design. This is

based on the conceptual equivalence of nonlinear isolated response with that of an

equivalent elastic system undergoing harmonic oscillations. In this respect, this approach

is inherently not entirely realistic. Provisions for more accurate nonlinear time-history

methods are also included in the Guide Specifications, but not required for all systems.

Further, unlike the building code, the Guide Specifications implicitly suggest that some

measure of substructure yielding is also allowed in an isolated bridge system. However, no

guidance is given on how to account for the effect of this substructure yielding into the

procedure.

6

Although the Guide Specifications are a major advance in bridge engineering, it is

apparent that the limited assessment of the response of complete, seismically isolated

bridge systems through analytical studies, laboratory tests or exposure to actual

earthquakes, and the variety of unresolved concerns have impeded the application of

protective systems technology to bridges. As such, research is needed directed towards

developing a better understanding of the response of seismically isolated bridge structures

and the adequacy of currrent design criteria.

1.3 The Protective Systems Research Program

1.3.1 Overall Objectives of the Program

In consideration of these research needs, a series of studies have been undertaken under

the Protective Systems Research Program sponsored by the California Department of

Transportation. A variety of studies were developed as part of this program in order to:

1. Evaluate the effect of bi-directional loading on seismic isolation bearings and developimproved analytical bearing models.

2. Establish an understanding of global and localized response characteristics of simpleseismically protected bridges subjected to various types of seismic input including far-field, near-fault, and soft-soil motions.

3. Validate the efficacy of seismic protective systems for bridges of more complex(realistic) configurations.

4. Assess the applicability of requirements in the AASHTO Guide Specifications forSeismic Isolation Design and recommend improvements based on the results of theresearch.

1.3.2 Overall Tasks Undertaken

Seven major tasks were undertaken to meet the overall objectives of this program. The

tasks were organized to improve knowledge of component behavior and use component

information to understand bridge system performance through integrated experimental

7

and analytical studies. The experimental studies of seismic isolation components and

model testing of seismically protected bridges made extensive use of the earthquake

simulator and bearing test machines at the Pacific Earthquake Engineering Research

Center (PEER), University of California, Berkeley. The major tasks in the research

program were as follows.

1. Experimentally characterize the properties of representative elastomeric and slidingisolation bearings under a wide variety of uni-directional and bi-directionaldisplacement histories and rates.

2. Use the experimental force-displacement data for isolator bearings to develop andcalibrate mathematical models for elastomeric and sliding isolation bearings suitablefor nonlinear response-history analysis of complete bridge systems.

3. Investigate analytically and experimentally the response of a simple, isolated bridgesystem to uni-directional and bi-directional earthquake-history inputs, representingfar-field and near-fault earthquake shaking on different soil types.

4. By experiment and analysis, investigate the effectiveness of fluid viscous dampers onthe earthquake response of simple bridges, including consideration of the effect ofdamper configuration.

5. Use experimentally validated analytical models to examine the effect of substructurestiffness, substructure mass, substructure damping and varying isolation systemproperties on the response of seismically isolated bridge overcrossings subjected tovarious seismic inputs.

6. Use a versatile shaking table model of a more complex bridge with one and two spans,supported on flexible piers, to investigate the effect of mass and stiffness eccentricity,pier flexibility, pier mass and strength, and varying isolator properties on the bridgesystem subjected to one to three components of earthquake ground motionrepresenting far-field, near-fault, and soft-soil earthquake shaking.

7. Evaluate the efficacy and applicability of requirements in the AASHTO GuideSpecifications for Seismic Isolation Design and recommend improvements based onthe results of the research.

1.3.3 Protective system devices considered

Three types of seismic isolation bearings were studied in the overall research program.

These included lead-rubber (LR) and high-damping rubber (HDR) elastomeric bearings

8

and Friction Pendulum (FP) sliding isolation bearings. At the time of writing, LR and FP

seismic isolation bearings have been used in bridge structures in the United States. Lead-

rubber and HDR bearings are composed of alternating layers of elastomer and steel

providing the bearing with sufficient stiffness to carry vertical loads. Energy dissipation is

provided by a central lead core in LR bearings and by a specially formulated elastomer in

the HDR bearing. Photographs of the LR and HDR bearings used in the program are

shown in Figure 1-2. The FP bearing is composed of a high axial load capacity, low-

friction composite coated slider riding on a spherical stainless steel surface. Energy

dissipation is achieved through friction as the articulated slider moves across the stainless

steel surface. An example of an installed FP bearing is shown in Figure 1-3. Detailed

specifications of these test bearings are provided in Chapter 3.

The fluid viscous damper has been the sole type of supplemental damping device used in a

bridge in California to this date. Fluid viscous dampers dissipate seismic energy as heat,

which is generated by the passage of the damper fluid (typically, a silicone oil) past or

through the damper piston. Figure 1-4 shows a typical cross-section through the type of

fluid viscous damper used in this study.

1.3.4 Test Models

In the Protective Systems Research Program, seismically isolated bridge systems were

investigated experimentally with two models on the earthquake simulator. The first model,

referred to as the rigid block model, is shown in Figure 1-5. This model represents a

seismically isolated bridge on rigid pier or abutment supports. The second model, referred

to as the elevated bridge model, is illustrated in Figure 1-6. Experimentation with this

9

latter model allowed testing of one- and two-span bridge configurations. Piers were

designed to provide a range of stiffness and strengths in the longitudinal and transverse

directions of the model. Pier stiffness could also be arranged to provide stiffness

eccentricity along the length of the bridge. Pier components also allowed substructure

mass to be varied. The model allowed dynamic testing with one, two, and three

components of ground motion.

1.4 Scope of this Report

This report addresses several key components of the above outlined Protective Systems

Research Program. The scope of this report relates to aspects of items 2, 3, and 4 of the

defined program objectives as outlined in Section 1.3.1, excluding evaluations related to

the use of supplemental damping devices. Specifically, Task 5, 6 and 7 of the program, as

outlined in Section 1.3.2, are performed herein. Analytical studies related to these tasks

utilize the mathematical models developed and calibrated under program Tasks 1 through

3 (see Section 1.3.2). The specific scope of this report is defined as follows:

1. System Studies

Investigate the response of simple, isolated bridge overcrossings to a variety of earthquake

time-history inputs, by analysis and experiment. Factors considered include the effect of

mass and stiffness eccentricity, pier flexibility, pier mass and strength, substructure

damping and isolator properties.

A scale model of a bridge with one- and two-span configurations on flexible piers was

designed and constructed for testing on the earthquake simulator. These experiments

10

facilitate validation of analytical results and provide proof-of-concept of the efficacy of

seismic isolation for various bridge configurations and ground motion inputs.

2. Design Methodology Studies

The applicability of several requirements in the AASHTO Guide Specifications for

Seismic Isolation Design are evaluated. Specifically, the efficacy of the Uniform Load

Method is investigated as it relates to the design of a broad range of seismically isolated

bridge systems. Also, the effects of bi-directional ground motions and substructure

yielding on the procedures are studied.

An extensive parametric study is performed to investigate the response of simple isolated

bridge overcrossings to a database of recorded earthquake time-history inputs.

Specifically, the effects of varying isolator properties, pier flexibility, pier mass,

substructure damping, and substructure yielding on system displacement and force

response are examined.

3. Conclusions and Recommendations

Overall conclusions of the research findings are identified. Recommendations are

developed for implementation of findings and/or improvement of current design

specifications as indicated by the research results.

1.5 Organization of this Report

Chapter 2 contains a summary of key aspects of current design procedures for seismically

isolated bridges. Various system configurations used for short overcrossings and viaducts

are examined to develop a representative conceptual model of an isolated bridge system.

11

Preliminary analytical evaluations of the effect of basic system variations on isolated

bridge response are presented to help identify key parameters and priorities for subsequent

investigations.

Chapter 3 presents the development and results of the experimental studies undertaken.

Ten configurations of a bridge model are constructed to investigate the effect on system

response of mass and stiffness eccentricity, pier flexibility, mass and strength, and various

isolation system configurations. The models were subjected to various combinations of

one and two components of horizontal ground motion along with vertical ground motion

for earthquakes representing far-field, near-fault, and soft-soil characteristics.

Chapter 4 presents evaluations of the Chapter 3 experimental test results. These

evaluations are organized around the key issues identified in Chapter 2, where applicable.

Chapter 5 presents the results of several related analytical studies. These studies assess the

applicability of several design procedures contained in the AASHTO Guide

Specifications. In particular, the efficacy of the Uniform Load Method and the effects of

bi-directional motions and substructure yielding on the Guide Specifications procedures

are examined. The sensitivity of the response of simple isolated bridge overcrossings to a

database of recorded earthquake time-history inputs is then studied parametrically. The

effect of varying isolator properties, pier flexibility, pier mass, and substructure damping

on system response is also examined.

12

Chapter 6 offers overall conclusions, recommendations, and the needs for further research.

Chapter 6 summarizes the findings of the research in the context of assessing and

recommending improvements where needed to the AASHTO Guide Specifications.

Appendices provide design drawings of the elevated bridge test specimen with photo logs

of construction sequences and model details, result summaries from specimen pre-test

simulation analyses, datalogs for experimental test sequences, earthquake simulation

(shake table) test result summaries, and detailed summaries for each of the analytical

formulations developed in Chapter 5.

a. Pot bearing b. Curved sliding bearing

Figure 1-1 Traditional (non-seismic) bridge bearings

13

a. LR bearing b. HDR bearing

Figure 1-2 Elastomeric seismic isolation bearings

Figure 1-3 Friction pendulum seismic isolation bearing

14

Figure 1-4 Cross-section through a fluid viscous damper (courtesy of Taylor Devices)

Figure 1-5 Rigid-block bridge model on earthquake simulator at the Pacific Earthquake Engineering Research Center, University of California, Berkeley.

15

Figure 1-6 Elevated bridge model with flexible piers on earthquake simulator at the Pa-cific Earthquake Engineering Research Center, University of California, Berkeley.

16

2 Preliminary Evaluations

2.1 Introduction

The preliminary evaluations presented in this chapter provide an overview of design

procedures, standard analysis methodology, and the effects of system configuration on the

response of simple isolated bridge systems. It is the purpose of this chapter to summarize

these issues and effects as a foundation for the more in depth research and evaluation

undertaken in subsequent chapters. This overview outlines many of the critical issues

related to the application of protective systems to simple bridge overcrossings. Many of

these issues, as mentioned in Chapter 1, have impeded the application of this technology

to bridges. As such, these issues constitute key areas of needed evaluation.

2.2 AASHTO Guide Specifications

The AASHTO Guide Specification for Seismic Isolation Design [AASHTO, 1999] was

developed based on information available at the time of its preparation for buildings and

bridges, and adapts essentially the same equivalent linear procedure for isolation design

contained in the building code [U.B.C., 1994]. The dynamic characteristics of isolated

bridges, however, have several fundamental differences from those of isolated buildings.

For example, the flexibility and mass of tall, substructure piers below the isolators in a

bridge may cause significant force or displacement redistributions, altering the

17

effectiveness of the isolation system. Relatively rigid foundations in seismically isolated

buildings limit these concerns. Higher modes contributed by in-plane deck flexibility, or

interactions between several bridge components separated along expansion joints, may

also amplify local component demands in isolated bridge systems, unlike more compact

and regularly shaped buildings. In this respect, it is not certain whether the procedures in

the Guide Specifications adequately consider these and other differences.

The Guide Specifications essentially provide a linearized procedure for isolated bridge

design. The Uniform Load Method prescribed in this document is the linearized procedure

for estimating the design response (i.e., peak forces and deformations) along each of the

orthogonal axes of an isolated bridge. This procedure assumes an equivalence of isolated

response with that of an elastic substitute system oscillating harmonically. The Guide

Specifications stipulates linearized properties based upon the overall isolated system’s

secant stiffness defined at maximum displacement. These properties rely on the

assumption that the isolated system may be represented by an idealized single-degree of

freedom “substitute” system for purposes of determining the global response of the bridge.

For this representation, “equivalent” damping properties are postulated by equating the

energy dissipated per cycle in the isolated bridge system in this state to that of an

“equivalent” linear visco-elastic damping component (oscillating harmonically at the

same maximum amplitude and a harmonic frequency characterized by the system mass

and secant stiffness). Design displacement is then estimated using a smoothed design

spectra modified to account for this “equivalent damping.” Because these linearized

properties are based upon the design response, this method is iterative.

18

Several concerns have been raised about this linearization approach. First of all, it has

been shown that linearized methods based on ideal harmonic response are not as accurate

as those based upon random response [Iwan, W.D., and Gates, N.C., 1979], or for that

matter, as explicit nonlinear methods. Provisions for more accurate time-history methods

are stipulated in the Guide Specifications, but are only required for isolation systems

exhibiting relatively large “effective” damping (in excess of 30 percent of critical). Thus, a

degree of inaccuracy is associated in the current formulation. Further, the idealized single-

degree-of-freedom representation utilized in the formulation makes accounting for

substructure mass or damping difficult, and the Guide Specifications provide no guidance

on how or when to consider these contributions. Simplified methods for accounting for

these effects have been presented [Sheng et al., 1994]. However, a multi-degree-of-

freedom structural representation may be required in certain cases to sufficiently account

for these higher mode contributions to local component responses.

The Guide Specifications stipulate that system response is be determined, utilizing these

simplified procedures and modeling assumptions, separately along each of the specified

orthogonal axes of the isolated bridge system. The Guide Specifications specifies that bi-

directional peak force response may be determined for design utilizing the familiar 100

percent plus 30 percent rule for combination of uni-directional maxima. The reliability of

this procedure, which is based on the results of previous random vibration studies on

reinforced concrete hysteretic systems [e.g., Park et al., 1986], is uncertain for application

to nonlinear isolated bridges. More accurate bi-directional and three-dimensional time-

history methods may be used, as stipulated in the Guide Specifications, but again these are

only required for certain systems.

19

Unlike the building code formulation, there is also an implicit assumption in the

formulation of the Guide Specifications that some measure of substructure yielding will

occur during the seismic response of an isolated bridge. This is inferred from the wording

of the Guide Specifications commentary which states that the lower response modification

factors (R-Factors) utilized for isolation design “ensure, on the average, essentially elastic

substructure behavior in the design-basis earthquake.” This stipulation of essentially

elastic behavior “on average” suggests inelastic substructure behavior will occur for

response above the mean. No further guidance is given, however, on how to incorporate

the effect of this substructure yielding into the procedure.

It is apparent that additional studies are needed to assess and, where necessary, improve

design criteria for seismically isolated bridge systems. It is critical to establish an

understanding of the unique features which effect the dynamic response of simple isolated

bridge overcrossings as a rational basis for this assessment. To this end, several

preliminary analytical studies are presented in the following section to help identify key

parameters to be emphasized through the remainder of this report.

2.3 Parametric Seismic Analysis of Idealized Elevated Isolated Bridges

2.3.1 Introduction

The nonlinear response of simple seismically isolated bridge systems, such as those shown

in Figure 2-1, is sensitive to many factors. These factors include but are not limited to: (1)

the characteristics of the earthquake input; (2) isolation device characteristics; (3)

substructure flexibility; (4) substructure mass; (5) higher modes; (6) system eccentricity;

and (7) substructure yielding.

20

A pilot parametric analysis was undertaken to assess the sensitivity of simple isolated

bridge systems to changes in ground motion characteristics and several of these key

system factors. These studies were performed using nonlinear response-history analysis

procedures and numerical models developed specifically for this undertaking.

2.3.2 Idealized Modeling Assumptions

A typical structural model of an isolated bridge bent responding uni-directionally in its

isolated mode is presented schematically in Figure 2-1. For purposes of design, the Guide

Specifications allow an idealized bilinear force-deformation relationship to be utilized to

characterize isolation bearing hysteresis and total system uni-directional response (for a

isolated bridge responding laterally in its isolated mode along either principal orthogonal

axis), as shown in Figure 2-2(a) and Figure 2-2(b), respectively. With this idealization,

three independent fundamental parameters define the overall bilinear system hysteresis.

These parameters could be chosen as the systems’ first- and second- slope stiffness (K1 &

K2, respectively) and the yield strength (Fy).

Utilizing these definitions (see Figure 2-2), the isolated bridge system properties may be

defined for convenience by the following parameters: the isolator characteristic strength

coefficient,

(1)

the isolator second-slope period,

(2)

CyisoFyMg--------=

Tiso 2π M kd⁄=

21

and the characteristic substructure period,

(3)

For an assumption of rigid-plastic isolation bearings (i.e., infinite ku), variations in the

isolator characteristic strength coefficient, isolator second-slope period, and the

characteristic substructure period effect variations in each of the three fundamental system

hysteretic parameters. In the pilot analytical studies isolation bearings were, for simplicity,

considered rigid-plastic in all cases and response to variations in hysteretic system

properties were examined parametrically by varying the three system properties defined in

Equation 1, 2, and 3 above.

Figure 2-1 Structural idealization of an isolated bridge bent

a. Hysteresis of isolation bearing b. Hysteresis of isolated bridge system

Figure 2-2 Bilinear idealization of an isolation bearing and an isolated bridge system

Tsub 2π M Ksub⁄=

K sub

M

di

d

keff

keff= effective isolator stiffnessKsub= linear substructure stiffnessM = lumped deck massdi = isolator displacementdsub = substructure displacementd = total displacement = di + dsub

ered in these studies assuming a bilin

2-3(b). With this characterization,

system properties (i.e., first- & seco

l t F ) b l t d i

idealized behavior of a subject isola

-1). Parameters are specified to match

e Specifications as shown. The fo

sentation of the subject isolated bri

idge overcrossing

Ksub1 and ku ∞

with rigid abutments and rigid-plastic isolators

spectively. This model is similar to

ch orthogonal axes of the bridge w

onent definitions of the AASHTO Gu

n response of the isolation componen

al bridge model

s

mass, stiffness, anddamping

representation (i.e., K1) may be compu

g components as

= kd and the isolation system defines

of a broad range of full scale (prototy

nstant with a characteristic yield stren

second-slope isolator period of Tiso

e 5 percent damped elastic displacem

represented.

t response data as presented in Figur

ubstructure flexibility representing Ts

e first- and second-slope stiffness of

e substructure becomes more flexible,

on 6) may be computed by combin

on of Equation 5 and the previous sys

ed in Table 2-2 were utilized. The tre

sistent for all ground motions conside

case.

r systems with the most rigid first-sl

ishes, in general, as T1 increases. T

ngth contours as the first-slope per

econds). For example, total displacem

ain considered. Substructure compon

periods of Tsub = 0.05, 1, and 2 seco

With this choice of system propert

ib i ld b d di

flexible substructure in this study,

percent to the total displacement for

ercent for the most flexible and weak

h h d ll i l i ib

olated response of symmetric bri

dies both the simplified single-degree

ridge analytical models were utilized

ions for highway bridges in the U.S.

Spans (ft) Σmpier/Mdeck

142, 100 0.04

124, 152, 124 0.17

70 0

mass becomes larger. These results

somewhat as the substructure becom

ucture becomes more flexible). This i

increasingly insignificant in contribu

ct of substructure mass contributions

rces. These evaluations are necessary

dered (or when they may be reasona

g ratio ζsub which acts on the total sys

with isolation bearings locked, as wo

n compute the viscous damping const

ture damping classically as

ve damping component, csub, assigned

erformed for representative substruc

Analyses were performed on full s

round motion histories selected from

ays greater than unity and tends to

ural components (i.e., Tsub approachin

ity, however, as substructure flexib

or example for this ground motion th

dies. The AASHTO Guide Specificati

each orthogonal axis of an isolated bri

ining uni-directional maxima with

n rule. To evaluate the efficacy of

o uni-directional ground motion inp

e peak total displacement response ve

results show that the 100 percent +

estimate of maximum bi-directio

ricity

y between the bridge superstructure m

his arrangement may be the produc

or unsymmetric skew for example

splacement response (d) for a system

ement response (d0) for the same sys

f these displacements, d/d0, increase

However the d/d0 ratio never excee

torsional response [Bozzo, et al, 19

stiffness and strength properties to a

to align with the center of superstruc

ke simulation studies may also provid

eneralized multi-degree-of-freedom,

d Section 2.3.3). Full scale (prototy

nge of substructure stiffness eccentri

center of mass and substructure cente

ut is limited to a moderate value

fault ground motion. This is an indica

s have on system torsional response e

a bridge of 100 foot span, this rota

), with peak rotation on the order of 0

ectively.

eak bridge deck rotation for an additio

rall isolated system. Balancing the in

blies along the bridge span by pro

o reduce torsional response (compare

ing at all supports). However, the lim

on

iffness eccentricity

can FP 2.0 1.0 DView FN 6.4 1.0 DView FP 6.4 1.0 DJMA FN 3.4 1.0 DJMA FP 3.4 1.0 Do, ivir 45 10 2.01 F

sponse vs. elastic response

0 0.25 0.5 0.75 1 1.25 1.5 1.75 20

2

4

6

8

10

12

14

Tsub

(sec)

Dis

plac

emen

t (in

)d d

i

elastic, 5% damping

3 3.1 3.2 3.3 3.4 3.5 3.66

8

10

12

T2 (sec)

Dis

plac

emen

t (in

)

d elastic: T

2, 15% damping

2 2.5 3 3.5 40

10

20

30

40

T2 (sec)

d (in

)

ope flexibility and characteristic streng

2 2.5 3 3.5 40

0.25

0.5

0.75

1

T2 (sec)

d (in

)

Cyiso

=.04

.08

.12

2 2.5 3 3.5 40

0.25

0.5

0.75

1

T2 (sec)

d (in

)

2 2.5 3 3.5 410

15

20

25

30

35

Tiso

(sec)

d i (in)

ope flexibility and characteristic streng]

2 2.5 3 3.5 40.9999

1

Tiso

(sec)

d i/d

Cyiso

=.04

.08

.12

2 2.2 2.4 2.6 2.8 3 3.2 3.4 3.6 3.8 40.3

0.4

0.5

0.6

0.7

0.8

0.9

1

Tiso

(sec)

d i/d

yiso .06, Tiso 3 sec, [prototype scale

0 0.5 1 1.5 20

5

10

15

20

25

30

Tsub

(sec)

d/d 0

damping ratio = 0.050.10 0.20

0 0.25 0.5 0.75 1 1.25 1.5 1.75 20.7

0.8

0.9

1

1.1

1.2

1.3

Tsub

(sec)

(dm

ax +

0.3

d min

)/d xy

LA01/02 LA21/22 NF01/02 LS17c/18c

n response, Cyiso .06, Tiso 3 sec, [pc/18c]

0 0.25 0.5 0.75 11

1.1

1.2

1.3

T1 (sec)

d/d0

e/L = .05.10

0 0.15 0.3 0.450

0.005

0.01

substructure stiffness eccentricity: e/L

peak

rot

atio

n (r

ad)

0 0.15 0.3 0.450

0.02

0.04

0.06

0.08

substructure stiffness eccentricity: e/L

peak

rot

atio

n (r

ad)

abutment Tiso

= 2 sec

3 sec 4 sec

eccentricity, [prototype scale, NF01]

67

3 Testing of a Bridge Deck Model

3.1 Introduction

The experimental studies in this research were developed to satisfy the program objectives

outlined in Section 1.3.1. Specifically, these studies fulfill program Task 6 in Section

1.3.2, to construct a scale model of a bridge with flexible piers for earthquake testing.

These tests utilized the newly upgraded 3-dimensional capabilities of the earthquake

simulator at the Pacific Earthquake Engineering Research (PEER) center and provided a

first-time opportunity to study the bi-directional and 3-dimensional response of isolated

bridge systems.

The preliminary analytical evaluations of Chapter 2 illustrate the basic sensitivity of the

response of isolated bridge overcrossings to various structural and isolator characteristics.

Based on these evaluations. the specimen tested were developed to confirm these

sensitivities, including the effect of mass and stiffness eccentricity, pier flexibility, pier

mass and strength, and varying isolator properties, on system response.

These simulation tests also subjected various isolated bridge overcrossing configurations

to multiple signals of one to three components of earthquake ground motion. These

earthquake signals were representative of significant far-field, near-fault, and soft-soil

seismic events. In this respect, these studies provide invaluable proof-of-concept

68

verification for the effectiveness and robustness of seismic isolation for simple bridge

types.

3.1.1 Similitude Requirements

Similitude requirements for dynamic structural models have been developed in previous

studies [e.g., Krawinkler, Moncarz, 1982]. These similitude constraints are defined by the

chosen length scale factor, 1r, prescribed for relating model size to the subject prototype

(full-scale) structure (i.e., lr = lm/lp, where lm and lp are the model and prototype length

scale, respectively). With gravity a constant and masses reduced proportionally by the

length factor, ground motion accelerations are held constant in dynamic testing to preserve

force scaling (where F = ma). Similitude requires that prototype and model periods be

related by the square root of the length scale factor (i.e., the time scale factor, tr = 1r1/2),

with frequencies related by its inverse (i.e, the frequency scale factor, ωr = lr-1/2). Ground

motion acceleration time-history records must then be “compressed” by multiplying their

time increment by the time scale factor to preserve similitude. This produces time-history

records for shake table simulations of equal acceleration output (as the original recorded

motion), but of higher frequency. With these constraints applied specimen simulations are

performed, with model and prototype system displacements (i.e., dm and dp, respectively)

related directly by the length scale factor (i.e., the displacement scale factor, dr = dm/dp =

1r) [Krawinkler, Moncarz, 1982].

In Chapter 2, isolated bridge behavior was shown in several cases to vary with system

properties corresponding to “short” and “long” spectral period ranges. The typical

response of these systems was shown in several cases to have unique characteristics in

69

each of these period ranges. To model this behavior experimentally, similitude requires

that the model period, Tm, be related to the prototype system period, Tp, by the time factor

(i.e., tr = 1r1/2) such that

(12)

With this requirement, model stiffnesses (Km) can be selected to produce required model

periods (i.e., Tm), given a model mass (Mm), through the definition of system period as

(13)

A length scale factor of 1/4 (i.e., lr = 1/4) was chosen for model scaling for these studies.

This ratio was selected based upon the load and geometric constraints of the earthquake

simulator which were reached for bridge model weight and length quantities represented

by a typical single-span bridge overcrossing reduced to this size. Ground motion scaling

was performed separately, with length scale factors chosen between 1/2 to 1/5 (i.e., lr

between 1/2 and 1/5) and related time scale factors between and (i.e., tr = 1/ and

1/ , see Equation 12), to produce ground motion simulation records of peak velocity and

displacement within the capacity of the shake table simulator (see Section 3.2 below).

The resulting relationships between model and prototype response produced by tests using

these separate model and ground motion scaling factors may be deduced by applying the

previous similitude relations. For example, for a given model with length scale factor of

trTmTp------- lr

1 2⁄==

Km 4π2MmTm--------=

2 5 2

5

70

lrm, a full-scale prototype system with a period of Tp = Tmlrm-1/2 is represented (see

Equation 12). Applying a different length scale factor of lrg to the ground motion implies

that this same model represents a prototype system with a period of Tp’ = Tmlrg-1/2

subjected to this scaled ground motion record (Equation 12). Combining these provides

Tp’ = Tp(lrm/lrg)1/2. Thus if the ground motion scale is smaller than the model scale (e.g.,

lrg = 1/5 and lrm= 1/4), testing with this ground motion represents the response of a

prototype system with a period, Tp’, larger than the full scale target prototype, Tp. On the

other hand, if the ground motion scale is larger than the model scale (e.g., lrg = 1/2 and

lrm= 1/4), the testing represents the response of a prototype system with a period, Tp’,

smaller than the full scale target prototype, Tp. In this way, the model (with fixed scale,

e.g., lrm = 1/4) may represent systems in different spectral regions of the ground motion

record by varying the chosen scale factor (i.e., lrg) applied to the motion in the given test

sequence. In either case, displacements to the full-scale motion are related to model

displacements in the test sequence by the ground motion scale factor, lrg, as before.

3.1.2 Design criteria

In order to illustrate the effect of basic system variations on the response of the isolated

bridge specimen, component properties were established to accommodate the same

substructure and isolation system variations as those evaluated in Chapter 2. These

properties were established to model systems over the broad spectral range where

preliminary analyses indicated significant variations in response.

71

Bridge size and mass were chosen to produce the largest similitude length scale within the

limits of the shake table physical geometry and vertical load capacity. This implied a ¼-

scale model (i.e., length scale factor, lr = 1/4) with maximum span of 20 feet and a

corresponding bridge deck mass of 65 kips.

Target isolation system properties were determined for the given bridge deck mass. It was

established that model isolation bearings with characteristic periods of Tiso = 1.3 to 1.75

seconds were desired (see Equation 2). These would represent full-scale (prototype)

isolated systems with a rigid-based isolation periods ranging from Tiso = 2.6 to 3.5 seconds

at lr = 1/4. Isolator strengths were selected to be in the typical range of Cyiso = 0.04 to 0.12

(see Equation 1). Several bearing types were to be represented to provide hysteretic

properties covering these ranges, including Lead-rubber (LR), High-damping rubber

(HDR), and Friction Pendulum (FP) bearing types.

Test specimen non-isolated periods of approximately Tsub = 0.025 to 1 seconds were

desired (see Equation 3). These represent prototype (full-scale) non-isolated bridges with

periods ranging from Tsub = 0.05 to 2 seconds at a length scale factor of lr = 1/4 which

covers the range of substructure systems evaluated in Chapter 2. The flexibility of

specimen substructure components was established to target substructure stiffness (Ksub)

to isolator second-slope stiffness (kd) ratios in the range of 2:1 to 7:1. As the AASHTO

Guide Spec presumes some measure of substructure yielding to occur (see Section 2.2), it

was also considered important to study its effect on the response of isolated bridge

systems. Therefore, weaker specimen substructure components designed to yield near

peak response were desired as part of the test specimen arrangement.

72

A versatile specimen which could accommodate these variations in pier flexibility and

isolation bearing properties was required for these studies. In addition, the model had to

accommodate varying pier mass and strength and eccentric arrangements of superstructure

mass and substructure stiffness. The design also needed to be easily alterable during

testing to accommodate these system changes efficiently.

3.1.3 Design Development

In keeping with these requirements, a ¼-scale bridge specimen design was developed with

preliminary consideration given to several alternate systems. Steel columns were

preferred because of the desire to perform many tests on the specimen without damage to

the total substructure. A simple design utilizing standard pipe column sections with

optional cross bracing (added for rigid or eccentric configurations) was first considered.

Standard pipe material (i.e., nominal yield strength, Fy = 46 ksi) proved inadequate to

resist combined effects of axial and flexural loads at pipe lengths required to achieve

target substructure flexibility. A variety of other ideas were then considered, including an

alternative utilizing steel wide-flange, L-frames mounted on 3-dimensional clevis base

pins. For this alternative, prestressed coil spring assemblies were designed to provide

variable rotational base flexibility. These assemblies allowed variations in the lateral

frame stiffness to be achieved by replacing springs of pre-engineered stiffness in the base

connection assembly. Connection detailing, necessary to ensure smooth movement of the

assemblies through the anticipated displacement ranges, proved complicated, however,

causing this alternative to be abandoned in lieu of the simpler final specimen design. Final

specimen design was taken as a combination of these two alternatives, as described in the

next section.

73

3.1.4 Final Specimen Design

Complete construction plans of the test specimen, including as-built photographs, are

included in Appendix A for reference. The final specimen substructure design utilized

steel tube frames mounted on clevis base pins. The clevis pin mounts rotate about two-

orthogonal axes and provide adequate transverse and longitudinal rotation capacity at the

base (with restraint about the vertical axis). Rotational stiffness was provided by thin

tapered plates extending in each orthogonal direction from the base of the steel tube

frames. The wide end of these plates attach to the tube frames near the clevis pins while

their narrow tapered ends attach to a shear pin connection which provides vertical support

only. Tube columns support load cells and isolation bearings, which in turn support a rigid

steel deck frame spanning between pier assemblies. The tube column frames support

vertical gravity load (imposed by the 65 kips of concrete blocks mounted on the bridge

deck frame as mass) and provide lateral stiffness supporting shear loads transmitted from

the bridge deck through the isolation bearings during seismic excitation.

Two tapered plate designs were prepared for the elevated bridge specimen. These two

designs allowed substructure lateral stiffness to be altered by changing the rotational

flexibility at the base of the piers (see Appendix A). Short and long tapered plates

provided the specimen with non-isolated periods of approximately Tsub = 0.67 and 1

seconds, respectively, Adding cross-bracing between the columns provided the specimen

with a non-isolated period of approximately Tsub = 0.25 seconds. Two sets of each size

tapered plates were fabricated from A514 and A36 steel, respectively. The stronger A514

material provided essentially elastic stress levels, while the weaker A36 material allowed

moderately ductile substructure behavior, during the test sequences.

74

Three separate sets of test bearings were designed and fabricated for use in the bridge deck

model tests: 1) high-damping rubber (HDR) bearings supplied by Bridgestone, Inc.; 2)

lead-rubber (LR) bearings supplied by DIS, Inc.; and 3) spherical sliding, or Friction-

pendulum (FP) bearings supplied by Earthquake Protection Systems, Incorporated.

Design details of each are included in Appendix A for reference.

The basic force diplacement properties of these three bearing types (see Figure 3.1) were

intentionally selected to be different in order to cover the ranges of isolation bearing

properties normally seen in practice and to examine behavior in different ranges of

behavior identified in the preliminary evaluations (see Chapter 2). One should not use

these tests to compare the effectiveness of different isolator types, as different designs

could also have resulted in isolators with very similar hysteretic properties. Results from

these tests should be used only for calibration of models, validation of identified trends

and behavior concepts (i.e., how system behavior is effected by changes in the

fundamental hysteretic properties Qd, ku and kd), and the efficacy of isolation in general.

Bearing hysteretic properties established from characterization test data are presented in

Table 3-1 as a reference. Figure 3-1 shows a uni-directional plot comparing the hysteretic

response of the HDR, LR, and FP test bearings illustrating the hysteretic differences in

strength and first- and second-slope stiffness characteristics of these bearing designs.

Characteristic target (rigid-based) isolation periods (i.e., Tiso, see Equation 2) are also

shown in Table 3-1 for each bearing type.

As illustrated in the plans (see Appendix A), this final specimen design is quite versatile.

Lateral substructure stiffness can be distributed uniformly at each span or

75

unsymmetrically (with increasing flexibility along the bridge length) to model bridges

spanning grade changes. Substructure flexibility can also be varied in the longitudinal and

transverse directions to model the cantilever versus frame behavior of a multi-column

bent. For two-span (six-pier) configurations, end piers can be braced longitudinally and

transversely and the central piers left relatively flexible to model bridge overcrossings.

Table 3-1 Summary of bearing mechanical properties from characterization tests

Bearing Type Qd (/W) ku(kips/in) kd (kips/in) Tiso4 (sec)

HDR1 0.015-0.05 4.5-3.5 1.75-1 1.29

LR2 0.05-0.09 15-5 1.6-0.9 1.36

FP3 µ = 0.07 - 0.10 INF R = 30 in.W/R= 0.54 kip/in 1.75

1. Property range over 50-250% shear strain with Tiso at γ = 250%2. Property range over 50-200% shear strain with Tiso at γ = 200%3. FP hysteretic properties based on friction coefficient, µ, and dish radius, R4. Tiso computed for a deck weight of 65 kips applied uniformly to four bearings

Figure 3-1 Comparison of test bearing hysteretic properties

0

0.05

0.1

0.15

0.2

0.25

0 0.5 1 1.5 2

d (in)

F/W

HDRLRFP

76

A column detail also allows additional substructure mass of up to 1.5 kips to be installed

near the 2/3rd height of each column frame. This represents substructure mass proportions

of up to γ = Σ mpier/Mdeck = 10 percent.

For some configurations, the isolators are mounted directly on the shake table. Separate

transition hardware allows FP, LR and HDR isolation bearings to be installed. Bridge deck

mass can also be arranged symmetrically or eccentrically. The mechanical properties of

the bridge specimen substructure are presented in Table 3-2.

Table 3-2 Summary of computed specimen substructure design properties

Frame Configuration

Braced1 Short Plate2 Long Plate3

Tsub 0.25 sec 0.67 sec 1.0 secKsub 26.6 kips/in 4 kips/in 1.8 kips/in

Fy N.A. 8 kips (A514)2.9 kips (A36)

8 kips (A514)2.9 kips (A36)

Cy = Fy/W N.A. 0.5 (A514)0.18 (A36)

0.5 (A514)0.18 (A36)

dy N.A. 1.9 in (A514)0.68 in (A36)

4.0 in (A514)1.44 in (A36)

Ksub/kd (LR) 27:1 4:1 2:1Ksub/kd (FP) 47:1 7:1 3:1

1. Braced pier assembly- see Figure A-1, A-2, & A-7 details, Appendix A2. Pier assembly with short plate - see Figure A-1, A-2, & A-10, Appendix A3. Pier assembly with long plate - see Figure A-1, A-2, & A-10, Appendix ATsub - see Chapter 2, Equation 3Ksub - lateral stiffness at top of pier assemblyFy, Cy - yield strength and yield strength coefficient of pier assemblydy - yield displacement at top of pier assemblyKsub/kd - substructure stiffness to isolator second-slope stiffness ratio

77

3.1.5 Test Set-Up and Protocol

Utilizing the bridge specimen substructure components and isolation test bearings

provided, ten bridge configurations were developed to study various aspects of the

response of isolated bridge overcrossings through earthquake simulation studies. This test

program was developed to examine the global and local response trends of simple and

more complex isolated bridge systems subjected to a variety of earthquake input,

fullfilling objective 2 and 3 of the research program (as outlined in Section 1.3.1). Specific

issues to be examined in the study were outlined in project Task 6 (see Section 1.3.2) and

discussed further in the preliminary evaluations of Chapter 2. These issues include the

effect on isolated bridge response of: (1) the characteristics of the earthquake input

(including fault proximity, site soil effects, and the effect of uni-directional vs. multi-

directional input); (2) isolation system characteristics; (3) substructure flexibility; (4)

substructure mass; (5) higher modes; (6) system eccentricity; and (7) substructure

yielding. The ten bridge specimen configurations developed for these studies are

described in the sections below and schematically illustrated in Figure 3-2 through 3-6.

Configuration 1

This configuration was the standard non-elevated symmetric mass configuration of the

bridge deck model (see Figure 3-2). Table motion in this case represents motion at “rigid”

abutments. In this configuration, load cells were mounted directly on the earthquake

simulator. Isolation bearings were mounted to the top of load cells and connected to the

underside of the steel bridge deck. Concrete blocks were prestressed to the deck frame,

providing 65 kips of weight. Instrumentation for this configuration included: (1) table

78

instrumentation measuring table accelerations and displacements; (2) five degree-of-

freedom load cells measuring x- & y- shear and moment, as well as axial forces, under

each isolation bearing (at quadrant Q1, Q2, Q3 and Q4); (3) linear potentiometers (LP)

measuring bearing, deck, and table displacements; (4) DCDT’s measuring potential

shifting of the concrete mass blocks ; and (5) accelerometers mounted to the bridge deck,

concrete mass, and locally above isolation bearings. Data acquisition in this configuration

is illustrated in Figure 3-7 and detailed descriptions of each instrumentation device are

outlined in Table 3-3.

Configuration 1 was developed to study the response of simple isolated bridge

overcrossings. Variations in isolator characteristic properties (i.e., first- and second- slope

stiffness and strength) were examined in this configuration by replacing bearing types.

The effect of earthquake motion characteristics, including bi-directional and vertical input,

fault-proximity and soil effects, were studied in this and all model configurations. The

effect of idealized near-fault pulse motions were studied solely in this basic configuration

(see Section 3.2 for information regarding earthquake test motions).

Configuration 2

This configuration was similar to Configuration 1, but incorporated an eccentric mass

layout for the bridge deck (see Figure 3-2). This configuration was erected similar to

Configuration 1 with concrete blocks offset to produce superstructure mass eccentricities.

Instrumentation for this configuration was the same as that for Configuration 1 (see Figure

3-7 and Table 3-3).

79

Configuration 2 was developed to study the torsional response of simple isolated bridge

overcrossings. Deck mass was systematically offset to produce superstructure mass

eccentricities of 5 and 10 percent of the longitudinal bridge span. Torsional response of

different isolation systems were also examined in this configuration by replacing bearing

types.

Configuration 3

This configuration was again similar to Configuration 1, but incorporated an unsymmetric

distribution of isolation bearings (see Figure 3-3). This configuration was erected similar

to Configuration 1 with two different isolation bearing types (HDR and LR bearings)

mounted at either end of the bridge span. Instrumentation for this configuration was the

same as that for Configuration 1 (see Figure 3-7 and Table 3-3).

Configuration 3 was developed to study the torsional response of isolated bridge

overcrossings produced by bearing stiffness eccentricity. The installation of HDR and LR

bearings at either end of the deck produced eccentricity in first- and second- slope stiffness

and bearing strength properties. This study was not performed to suggest that different

types of isolators could be installed in a bridge at opposite abutments, but to assess the

effect of these types of variations along a bridge span. The effect of earthquake input

characteristics on this stiffness induced torsional response were also studied in this

configuration.

80

Configuration 4

This was the standard four-pier elevated configuration of the bridge deck model, an

elevated single-span configuration with symmetric substructure stiffness (see Figure 3-3).

In this configuration, substructure pier assemblies were mounted directly to the

earthquake simulator with load cells mounted on top of pier supports. FP isolation

bearings were used exclusively in this configuration attached to the top of load cells and

connected to the underside of the steel bridge deck. Concrete blocks were prestressed to

the deck frame, providing 65 kips of weight. In addition to the same instrumentation

utilized for Configuration 1 through 3 (see Figure 3-7 and Table 3-3), the following

instrumentation was added in this configuration: (1) additional linear potentiometers (LP)

measuring top of pier displacements; (2) additional accelerometers mounted to top of

substructure piers; and (3) linear strain gauges mounted on tapered leaf springs at base of

piers to monitor potential yield conditions. Data acquisition hardware in this configuration

is illustrated in Figure 3-7, 3-8, and 3-9. Detailed descriptions of each instrumentation

device are outlined in Table 3-3 and 3-4.


overcrossings and viaducts. Variations in isolator characteristic strength were examined in

this configuration by replacing FP bearing slider types. The effect of variations in

substructure flexibility were examined by adding or removing bracing or using different

rotational leaf springs at the base of the pier assemblies.

81

Configuration 5

This was an elevated bridge configuration with un-symmetric substructure stiffness,

erected similarly to Configuration 4, with short and long leaf springs inserted at the base

of pier assemblies in the x- and y- directions, respectively (see Figure 3-4).

Instrumentation for this configuration was similar to Configuration 4 as illustrated in

Figure 3-7, 3-8, and 3-9 and outlined in Table 3-3 and 3-4.


overcrossings, with the effect of variations in substructure flexibility in the x- and y-

directions. Variations in isolator characteristic strength were examined in this

configuration by replacing FP bearing slider types as before. The effect of substructure

mass (where γ = Σ mpier/Mdeck) was also studied in this configuration by varying the

number of attached lead weight packets mounted to substructure pier assemblies.

Configuration 6

This configuration is similar to Configuration 5 with additional cross-bracing added to

piers at one end of the specimen (see Figure 3-4). Instrumentation for this configuration

was similar to Configuration 4 and 5 as illustrated in Figure 3-7, 3-8, and 3-9 and outlined

in Table 3-3 and 3-4.


overcrossings having unsymetric substructure conditions. Variation of substructure

flexibility along the span produced by added cross-bracing created the effect of

eccentricity in substructure stiffness in the configuration. Variations in isolator

82

characteristic strength could be examined in this configuration by replacing FP bearing

slider types symmetrically (i.e., replacing all slider types at once) or unsymmetrically (i.e.,

placing weaker FP bearing sliders above braced piers to counteract torsional response).

Configuration 7

This was the standard elevated double-span configuration of the bridge deck model (see

Figure 3-5). In this configuration, six substructure pier assemblies were mounted directly

to the earthquake simulator. Cross-bracing was added to piers in the longitude and

transverse directions at each end of the bridge specimen to simulate stiff abutment

conditions. Center piers were configured with short and long leaf springs inserted at their

base in the x- and y- directions, respectively, simulating a flexible central pier bent. Load

cells were mounted on top of pier supports with FP isolation bearings attached to the top

of load cells and connected to the underside of the steel bridge deck. Concrete blocks and

additional lead packets were prestressed to the deck frame providing 99 kips of weight.

This additional mass was installed to provide the system with an overall yield strength

coefficient (i.e., Cyiso) theoretically equivalent to the four pier specimen (assuming a

constant FP slider µ value). In addition, this larger mass produced nearly equivalent period

characteristics for the six-pier configuration with out cross-bracing as the four pier un-

braced configurations. The instrumentation for this configuration was similar to

Configuration 4 through 6, with the following additions: (1) five degree-of-freedom load

cells under each isolation bearing at pier 5 and 6; (2) linear potentiometers (LP) measuring

top of pier 5 and 6 displacements; (3) accelerometers mounted to locally above isolation

bearings and at top of substructure piers at pier 5 and 6; and (4) linear strain gauges

83

mounted on tapered leaf springs at base of pier 5 and 6. Data acquisition hardware is

illustrated in Figure 3-7, 3-8, and 3-9 and outlined in Table 3-3 and 3-4, as before.

Configuration 7 was developed to study the response of simple double-span isolated

bridge overcrossings. Variations in isolator characteristic strength were examined in this

configuration by replacing FP bearing slider types. These strength variations were

implemented symmetrically (i.e., by replacing all sliders at once) or unsymmetrically (by

placing weaker FP sliders selectively above braced piers or central piers to effect force

distribution).

Configuration 8

This configuration was similar to Configuration 7 with cross-bracing added to piers at

only one end of the specimen and long tapered leaf spring plates installed at the base of

piers in the x- direction at the opposite end (see Figure 3-5). Instrumentation for this

configuration was similar to Configuration 7 as illustrated in Figure 3-7, 3-8, and 3-9 and

outlined in Table 3-3 and 3-4.


overcrossings. Increasing substructure transverse flexibility along the span produced

eccentricity in substructure stiffness. The effect of varying isolator characteristic strength

along the span was examined by replacing FP slider types symmetrically (i.e., replacing

all slider types at once) or unsymmetrically (i.e., placing weaker FP bearing sliders above

braced piers to counteract torsional response).

84

Configuration 9

This configuration was similar to Configuration 7 with weaker (A36) tapered leaf spring

plates inserted at the base of central piers (see Figure 3-6). Instrumentation for this

configuration is similar to Configuration 7 as illustrated in Figure 3-7, 3-8, and 3-8 and



bridge overcrossings with yielding substructure components. A36 leaf spring plates in

central piers were designed to allow ductile response in the lateral force-displacement

characteristics of the central pier assemblies.

Configuration 10

This was a free-standing configuration of substructure pier sub-assemblies (see Figure 3-

6). This configuration utilized the same substructure assemblages as Configuration 9 with

bridge mass and deck, isolation bearings, and load cells removed. Instrumentation for this

configuration consisted of: (1) a linear load cell attached in series with a come-along

winch between tops of pier sub-assemblies to produce an increasing static force, and (2)

linear potentiometers (LP) measuring the displacements at the top of the piers (see Figure

3-6). This configuration was developed to characterize the lateral force-deformation

characteristics of the strong and weak substructure pier sub-assemblies (i.e., those

utilizing A514 and A36 tapered leaf spring base plates, respectively). The piers were

characterized in both their stiff (i.e., short plate) and flexible (i.e., long plate)

configurations (see Table 3-2).

85

Specimen as-built documentation

As a reference, photographs of bridge specimen erection sequences, pier frame

assemblies, and component details are included along with construction plans and

fabrication details in Appendix A.

3.2 Earthquake Histories for Testing

An evaluation of the effect of ground motion input characteristics on the response of the

various configurations of the bridge deck specimen was an essential component in these

simulation studies. The features of seismic demand include: (1) fault proximity (i.e., far-

vs. near- field events); (2) directivity (i.e., fault-normal vs. fault-parallel and forward- vs.

backward-azimuth motions); (3) site specific soil conditions (i.e., rock vs. soil sites); (4)

bi-directional effects; and (5) vertical components of motion. A group of shake table

motions was tailored to consider each of these variations explicitly. In addition, sinusoidal

forcing motions were developed to characterize the friction coefficient of the various FP

bearing slider types provided by the manufacturer. The characteristics of these motions are

described in detail below.

Three representative earthquake time-histories were selected for shake table testing. Basis

recorded time-histories were chosen from the database developed in Phase 2 of the

FEMA/SAC project [SAC, 1997]. The bi-directional ground motion pairs LA13/LA14,

NF01/NF02, and LS17c/LS18c, as listed in Table 2-2, were selected from this database to

represent far-field, near-fault, and soft-soil ground motion types, respectively. These basic

motions were filtered and scaled to produce shake table input motions conforming to the

acceleration, velocity, and displacement limits of the shake table apparatus. The resulting

86

Configuration 1: Non-elevated Symmetric Mass

Configuration 2: Non-elevated Eccentric Mass

Figure 3-2 Test Configurations 1 and 2 for bridge deck model

x

y

(4) FP bearingsor(4) LR bearingsor (4) HDR bearings

C.M.

Deck FrameWdeck = 66 kips

x

y

(4) FPS bearings

C.M.

ey

or(4) LR bearings


87

Configuration 3: Non-elevated Eccentric Stiffness

Configuration 4: Elevated Single-span with Symmetric Stiffness


x

y

(2) HDR bearings

(2) LR bearings


C.M.

x

y

(4) FPS bearings

C.M.

Deck frameWdeck = 66 kips

(4) Substructure piersγ=Σmpier/Mdeck = .05, typical

Stiff rotational springs in x & ydirections, typicalat base of piers

Q1Q2

Q3 Q4

~

88

Configuration 5: Elevated Single-span with Un-symmetric Stiffness

Configuration 6: Elevated Single-span with Eccentric Stiffness


x

y

(4) FP bearings

C.M.


(4) Substructure piers

Stiff rotational springs inx-direction, typ.

Flexible rotational springs iny-direction, typ.

Symmetric pier bracing in x & y direction, optional

γ=Σmpier/Mdeck = .05, typ.~ Single Span Elevated Deck Σmpier/Mdeck = 0.10 condition shown~

x

y

(4) FP bearings

C.M.


(4) Substructure piersγ=Σmpier/Mdeck = .05, typ.



Pier bracing added in x & ydirection, as shown

~

89

Configuration 7: Elevated Double-span

Configuration 8: Elevated Double-span with Eccentric Stiffness


(6) FP bearings





Symmetric pier bracing in x & y direction, typicalat end bays as shown

Q1Q2

Q3 Q4

Pier 5Pier 6

x

y

C.M.

~

(6) FP bearings


Stiff rotational springs at Q3 & Q4


Symmetric pier bracing in x & y direction at endshown only

Q1Q2

Q3 Q4

Pier 5Pier 6

x

y

C.M.

Flexible rotational springs at Q1 & Q2,x-direction


and Pier 5 & 6,x-direction

~

90

Configuration 9: Elevated Double-span with Yielding Piers 5 and 6

Configuration 10: Substructure Static Pullback Tests






Q1Q2

Q3 Q4

Pier 5Pier 6

x

y

C.M.

(6) FP bearings


Yielding rotationalspring plates inx & y-direction atPiers 5 & 6 only

(A514 - 100ksi)

(A514 - 100ksi nom.)

(A36 - 36ksi nom.)

~

Deck frame and bearings(6) Substructure piers



Q1Q2

Q3 Q4

56

x

y Yielding rotationalspring plates inx & y-direction atPiers 5 & 6 only(A36 - 36ksi nom.)

(A514 - 100ksi nom.)

(A514 - 100ksi nom.)

Typical pullback rigging:tension cable and straps,load cell, and come-alongin series at top of pier

removed, typical

91

Figure 3-7 Instrumentation plan for Configurations 1 through 3 of non-elevated bridge deck model

Q

◆ ◆ ◆

◆◆ ◆

◆◆

◆

92

Figure 3-8 Supplementary instrumentation for Configurations 4 through 9 of elevated bridge deck model

93

a. Short plate (a.k.a., stiff spring) b. Long plate (a.k.a., flexible spring)

Figure 3-9 Typical strain gauge layout at pier base rotational leaf springs

Table 3-3 Instrumentation for Configurations 1 thru 3 of non-elevated bridge deck

Channel No. Transducer Response

Quantity Orientation Transducer Location

1 LVDT displacement horizontal table h1o stroke2 LVDT displacement horizontal table h2o stroke3 LVDT displacement horizontal table h3o stroke4 LVDT displacement horizontal table h4o stroke5 DCDT displacement vertical table v1o stroke6 DCDT displacement vertical table v2o stroke7 DCDT displacement vertical table v3o stroke8 DCDT displacement vertical table v4o stroke9 A acceleration horizontal table h1-2 acc.10 A acceleration horizontal table h3-4 acc.11 A acceleration horizontal table h4-1 acc.12 A acceleration horizontal table h2-3 acc.13 A acceleration vertical table 1v acc.14 A acceleration vertical table 2v acc.15 A acceleration vertical table 3v acc.16 A acceleration vertical table 4v acc.17 LC force horiz. shear x load cell - bearing Q118 LC force horiz. shear y load cell - bearing Q119 LC force moment y load cell - bearing Q120 LC force moment x load cell - bearing Q121 LC force axial load cell - bearing Q122 LC force horiz. shear x load cell - bearing Q223 LC force horiz. shear y load cell - bearing Q224 LC force moment y load cell - bearing Q225 LC force moment x load cell - bearing Q2

C L

3”

linear strain gauge (SG),top & bottom of plate

Plan View (N.T.S.)

C L

3” 12”


SG, top ofplate only

Plan View (N.T.S.)

94

26 LC force axial load cell - bearing Q227 LC force horiz. shear x load cell - bearing Q328 LC force horiz. shear y load cell - bearing Q329 LC force moment y load cell - bearing Q330 LC force moment x load cell - bearing Q331 LC force axial load cell - bearing Q332 LC force horiz. shear x load cell - bearing Q433 LC force horiz. shear y load cell - bearing Q434 LC force moment y load cell - bearing Q435 LC force moment x load cell - bearing Q436 LC force axial load cell - bearing Q437 A acceleration horiz. y mass - acc138 A acceleration horiz. y mass- acc239 A acceleration horiz. x mass - acc340 A acceleration horiz. x mass - acc441 A acceleration horiz. x mass - acc542 A acceleration horiz. y deck frame - acc643 A acceleration horiz. y deck frame - acc744 A acceleration horiz. x deck frame - acc845 A acceleration horiz. x deck frame - acc946 A acceleration horiz. x deck frame - acc1047 A acceleration horiz. x bearing Q1 - acc1148 A acceleration horiz. y bearing Q1 - acc1249 A acceleration vertical bearing Q1 - acc1350 A acceleration horiz. x bearing Q2 - acc1451 A acceleration horiz. y bearing Q2 - acc1552 A acceleration vertical bearing Q2 - acc1653 A acceleration horiz. x bearing Q3 -acc1754 A acceleration horiz. y bearing Q3 - acc1855 A acceleration vertical bearing Q3 - acc1956 A acceleration horiz. x bearing Q4 - acc2057 A acceleration horiz. y bearing Q4 - acc2158 A acceleration vertical bearing Q4 - acc2259 LP displacement horiz. y deck frame - LP160 LP displacement horiz. y deck frame - LP261 LP displacement horiz. x deck frame - LP362 LP displacement horiz. x deck frame - LP463 LP displacement horiz. x deck frame - LP564 LP displacement horiz. y bearing Q1 - LP665 LP displacement horiz. x bearing Q1 - LP766 LP displacement horiz. y bearing Q2 - LP867 LP displacement horiz. x bearing Q2 - LP968 LP displacement horiz. y bearing Q3 - LP10


95

69 LP displacement horiz. x bearing Q3 - LP1170 LP displacement horiz. y bearing Q4 - LP1271 LP displacement horiz. x bearing Q4 - LP1372 LP displacement horiz. y table - LP1473 LP displacement horiz. x table - LP1574 LP displacement horiz. x table - LP16

75 DCDT displacement sloped deck frame to mass - DCDT1



Table 3-4 Instrumentation for Configurations 4 through 10 of elevated bridge deck



1 thru 58 |Same as Configurations 1 through 3 - non-elevated Bridge Deck, see Table 3-3 above

59 A acceleration horizontal Pier Q1 x - acc2360 A acceleration horizontal Pier Q1 y - acc2461 A acceleration horizontal Pier Q2 x - acc2562 A acceleration horizontal Pier Q2 y - acc2663 A acceleration horizontal Pier Q3 x - acc27

Channel 63 converted to uni-axial force readings for Config. 10 pullback tests, as follows63 LC force horizontal Uni-Axial load cell64 A acceleration horizontal Pier Q3 y - acc2865 A acceleration horizontal Pier Q4 x - acc2966 A acceleration horizontal Pier Q4 y - acc30




70 SG linear strain curvature y Pier Q1 - XOP71 SG linear strain horiz. x Pier Q1 - XFT72 SG linear strain horiz. x Pier Q1 - XTH73 SG linear strain curvature x Pier Q1 - YOP74 SG linear strain horiz. y Pier Q1 - YFT


96

75 SG linear strain horiz. y Pier Q1 - YTH76 SG linear strain curvature y Pier Q2- XOP77 SG linear strain horiz. x Pier Q2- XFT78 SG linear strain horiz. x Pier Q2 - XTH79 SG linear strain curvature x Pier Q2 - YOP80 SG linear strain horiz. y Pier Q2 - YFT81 SG linear strain horiz. y Pier Q2 - YTH82 SG linear strain curvature y Pier Q3 - XOP83 SG linear strain horiz. x Pier Q3 - XFT84 SG linear strain horiz. x Pier Q3 - XTH85 SG linear strain curvature x Pier Q3 - YOP86 SG linear strain horiz. y Pier Q3 - YFT87 SG linear strain horiz. y Pier Q3 - YTH88 SG linear strain curvature y Pier Q4- XOP89 SG linear strain horiz. x Pier Q4- XFT90 SG linear strain horiz. x Pier Q4 - XTH91 SG linear strain curvature x Pier Q4 - YOP92 SG linear strain horiz. y Pier Q4 - YFT93 SG linear strain horiz. y Pier Q4 - YTH94 SG curvature curvature y Pier 5 - XOP95 SG linear strain horiz. x Pier 5 - XFT96 SG linear strain horiz. x Pier 5 - XTH97 SG curvature curvature x Pier 5 - YOP98 SG linear strain horiz. y Pier 5 - YFT99 SG linear strain horiz. y Pier 5 - YTH100 SG curvature curvature y Pier 6- XOP101 SG linear strain horiz. x Pier 6- XFT102 SG linear strain horiz. x Pier 6 - XTH103 SG curvature curvature x Pier 6 - YOP104 SG linear strain horiz. y Pier 6 - YFT105 SG linear strain horiz. y Pier 6 - YTH106 LC force horiz. shear x load cell - bearing 5107 LC force horiz. shear y load cell - bearing 5108 LC force moment y load cell - bearing 5109 LC force moment x load cell - bearing 5110 LC force axial load cell - bearing 5111 LC force horiz. shear x load cell - bearing 6112 LC force horiz. shear y load cell - bearing 6113 LC force moment y load cell - bearing 6


97

114 LC force moment x load cell - bearing 6115 LC force axial load cell - bearing 6

Channels 116 through 128 skipped129 LP displacement horiz. y deck frame - LP1130 LP displacement horiz. y deck frame - LP2131 LP displacement horiz. x deck frame - LP3132 LP displacement horiz. x deck frame - LP4133 LP displacement horiz. x deck frame - LP5134 LP displacement horiz. y bearing Q1 - LP6135 LP displacement horiz. x bearing Q1 - LP7136 LP displacement horiz. y bearing Q2 - LP8137 LP displacement horiz. x bearing Q2 - LP9138 LP displacement horiz. y bearing Q3 - LP10139 LP displacement horiz. x bearing Q3 - LP11140 LP displacement horiz. y bearing Q4 - LP12141 LP displacement horiz. x bearing Q4 - LP13142 LP displacement horiz. y table - LP14143 LP displacement horiz. x table - LP15144 LP displacement horiz. x table - LP16145 LP displacement horiz. x Pier Q1 - LP17146 LP displacement horiz. x Pier Q1 - LP18147 LP displacement horiz. y Pier Q2 - LP19148 LP displacement horiz. x Pier Q2 - LP20149 LP displacement horiz. y Pier Q3 - LP21150 LP displacement horiz. x Pier Q3 - LP22151 LP displacement horiz. y Pier Q4 - LP23152 LP displacement horiz. x Pier Q4 - LP24153 LP displacement horiz. y bearing 5- LP25154 LP displacement horiz. x bearing 5 - LP26155 LP displacement horiz. y bearing 6 - LP27156 LP displacement horiz. x bearing 6 - LP28157 LP displacement horiz. y Pier 5 - LP29158 LP displacement horiz. x Pier 5 - LP30159 LP displacement horiz. y Pier 6 - LP31160 LP displacement horiz. x Pier 6 - LP32

The following Channels were relocated for Pier 5 & 6 acceleration readingsnote: acc6, acc7, acc8, acc10 to Pier 5x, Pier 5y, Pier 6x, Pier6y respectively

42 A acceleration horiz. x Pier 5 - acc31


98

table motions produced are referred to as LA13_14, NF01_02, and LS17c_18c and

utilized associated similitude length scale factors of lr = 1/2, 1/5, and 1/5, respectively, to

enforce conformance with table limits. Each table motion consists of a bi-directional pair

of time-histories oriented orthogonally along the x- and y- axes of the shake table (e.g.,

LA13_14 consists of the scaled x-direction LA13 and y-direction LA14 components). The

scaled near-fault motion (i.e., NF01_02) includes its vertical component as well. Motions

could be run separately, as uni-directional (x- or y-direction) components, or as bi-

directional pairs. In the case of the NF01_02 motion, the vertical component could be

added or omitted from the table time-history. Response spectra for the LA13_14,

NF01_02, and LS17c_18c shake table motions are shown in Figure 3-10 for comparison.

In addition, to study the effect of similitude scaling on the bridge specimen (see Section

3.1.1 discussion) these motions were processed with alternate length scale factors. The

effect of varying similitude length scale (lr) on the response spectra of three of these table

motion components is illustrated in Figure 3-11.

Near-fault ground motions have been shown to place significant demand on structures.

The impulsive and sometimes long-period content of these motions may produce velocity

and displacement demands significantly exceeding the design criteria provided by code

specified spectrum compatible ground motions. These unique earthquake time-histories

may be modeled with sufficient accuracy (for structures responding at or near the peak in

43 A acceleration horix. y Pier 5 - acc3244 A acceleration horiz. x Pier 5 - acc3346 A acceleration horix. y Pier 5 - acc34


99

the earthquake’s response spectra) as simple pulse motions [Krawinkler and Alavi, 1998].

On this basis, a suite of bi-directional, pure pulse motions were developed for these

studies with pure cosine and sine waves selected to model fault-normal and fault-parallel

displacement histories, respectively. Figure 3-12 illustrates time-histories for these

simulated motion pairs at their peak displacement amplitude of 5 inches (the limit of the

shake table apparatus). Fault-normal pulse durations of Tp = 1 and 2 seconds were selected

providing a range bracketing characteristic periods (i.e., Tiso) of the test bearings (see

Table 3-1). The duration of associated fault-parallel motions were taken as 2/3, 1, and 3/2

of Tp to evaluate the coupling effect between these orthogonal components. Table 3-5

shows amplitude and duration parameters for the entire suite of idealized pulse motions

developed from these parameters. These synthetic motions are representative of suggested

pulse parameters for actual near-fault earthquake records as recommended in [Krawinkler

and Alavi, 1998] at a similitude length scale factor of lr = 1/4.

To estimate friction coefficients for the various Friction Pendulum slider composites

provided by the manufacturer, a suite of sinusoidal acceleration time-histories were also

developed. These motions were designed to impose a steady-state response of varying

displacement and velocity amplitudes on the test specimen. Results from these tests were

utilized to estimate instantaneous friction values for the different composites. Figure 3-13

illustrates a representative acceleration time-history of these sinusoidal motions. Signal

amplitude and frequencies used to develop the various test motions are in Table 3-6.

100

a. Displacement spectra: LA13_14 b. Total acceleration spectra: LA13_14

c. Displacement spectra: NF01_02 d. Total acceleration spectra: NF01_02

e. Displacement spectra: LS17c_18c f. Total acceleration spectra: LS17c_18c

Figure 3-10 Displacement and total acceleration spectra for LA13_14, NF01_02, and LS17c_18c table motions, ζ=5%

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

la13la14

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

3

T (sec)

A to

tal (

g)

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

nf01nf02

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

3

T (sec)

A to

tal (

g)

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

ls17cls18c

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

3

T (sec)

A to

tal (

g)

101

a. Displacement spectra: LA13 b. Total acceleration spectra: LA13

c. Displacement spectra: NF01 d. Total acceleration spectra: NF01

e. Displacement spectra: LS17c f. Total acceleration spectra: LS17c

Figure 3-11 Response spectra of LA13, NF01, and LS17c table motions with varying length scale, ζ = 5%

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

SF=2SF=4

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

T (sec)

A to

tal (

g)

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

SF=3SF=5

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

T (sec)

A to

tal (

g)

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

SF=3SF=5

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

T (sec)

A to

tal (

g)

102

3.3 Pre-Test Analysis of Bridge Deck Model

An analysis matrix, developed to include all proposed model configurations (see Section

3.1.5), was prepared for pre-test analytical evaluations of the test specimen. Pre-test

analysis assumed properties for specimen substructure components based upon their

design values (see Section 3.1.4). For systems employing FP bearings, nominal friction

coefficients of 0.04, 0.06, 0.09, and 0.12 were considered to bound the range of expected

values. FP bearing stiffness was assumed based upon the design value of the dish radius, R

a. Fault-normal b. Fault-parallel

Figure 3-12 Idealized near-fault pulse displacement histories

Table 3-5 Characteristics of near-fault pulse table motions at full amplitude

Label Fault-orientation Tp (sec) Dgmax (in.) Vgmax (in./sec) Agmax (g)

nfpu

lse1

_

normal 1 5 31.4159 0.5108

parallel 0.667 5 47.1003 1.1483

parallel 1 5 31.4159 0.5108

parallel 1.5 5 20.9440 0.2270

nfpu

lse2

_

normal 2 5 15.7080 0.1277

parallel 1.33 5 23.6210 0.2888

parallel 2 5 15.7080 0.1277

parallel 3 5 10.4720 0.0568

0 0.5 1−5

0

5

time/Tp

Dg (

g)

0 0.5 1−5

0

5

time/Tp

Dg (

in)

103

(see Section 3.1.4). LR bearing properties were averaged from available characterization

test data. Analyses were performed utilizing the generalized multi-degree-of-freedom, bi-

directional bridge model developed previously (see Section 2.3.3). The three earthquake

time-history motions (i.e., LA13_14, NF01_02, and LS17c_18c) and the suite of idealized

near-fault pulse motions (i.e., nfpulse1_ and nfpulse2_, see Table 3-5) were considered

over a range of amplitudes.

Summarized results from these pre-test analyses are included in Appendix B. These

results were examined to establish table motion amplitudes which would likely produce

specimen response near the limits of bearing and component displacement capacities.

Figure 3-13 Representative acceleration time-history of sin signal table motion

Table 3-6 Sin signal characteristics for characterization of FP slider µ values

Ao (g) f (Hz) Do (in) Vo (in/sec)

.130 1.714 0.4331 4.6643

.145 2.285 0.2718 3.9024

.111 2.285 0.20807 2.98740

.097 2.856 0.11639 2.08867

.104 6.000 0.02827 1.0659

0 2 4 6 8 10 12 14−0.15

−0.1

−0.05

0

0.05

0.1

0.15

time (sec)

acc

ele

ratio

n (

g)

104

Trial span settings for test motions were then selected to provide 50% and 100% of these

peak amplitudes. Final peak table span settings were tailored during simulation tests to

produce model response near capacity, with final test sequences run at 50% and 100% of

these final settings.

3.4 Experimental Results for Configuration 1 (Non-elevated, Symmetric Mass)

Table C-1 in Appendix C presents a complete log of shake-table tests performed on the

Configuration 1 isolated bridge specimen. Selected experimental results from these tests

are presented in Appendix D as Figures D-1 through D-30. For each of the selected tests,

x-y orbit plots of the global displacement and global force, and x- & y-direction bearing

hysteresis (for quadrant 1, a.k.a., Q1) are presented as a summary of the test specimen

response. For uni-directional tests, x-direction displacement and force histories are

presented in lieu of orbits and y-direction bearing hysteresis at quadrant Q1 is omitted.

Selected results are shown only for the maximum table span setting run for each

individual test motion sequence.

The Configuration 1 specimen was tested utilizing the three selected earthquake table

motions and the suite of idealized near fault pulses (see Section 3.2). Tests were performed

at two or more amplitudes (or span settings) and bi-directional motion pairs were run

simultaneously and as separate x- & y-direction inputs. Sinusoidal bearing

characterization tests were also performed to estimate friction coefficients for FP bearing

slider composites.

For earthquake history tests, representative results for both uni-directional and bi-

directional tests are presented in Figures D-1 through D-19. These results are presented for

105

tests with FP, HDR, and LR bearings. Figure D-1 and D-5 illustrate response to the

LA13_14 earthquake table motion. Figure D-1 and D-2 compare response of a system

utilizing FP bearings with type 1 and 2 PTFE slider composite to the x-direction LA13

component. As seen, the FP bearings exhibit broad stable hysteretic response. The type 2

slider composite has a higher friction coefficient (see Section 3.14) resulting in a lower

peak displacement response. Figure D-3 shows the effect of the bi-directional LA13_14

table motion on the FP bearing system. As seen in the displacement orbit (see Figure D-

3(a)), this motion produces response oriented strongly along a 45 degree line to the xy-

axes. This behavior suggests strong coupling in the bearing response, as interaction with

the bearing yield surface is encountered. As expected, the x-direction bearing response at

quadrant Q1 (see Figure D-3(c)) is effected by this coupling, as the force response of the

bearing is reduced (compared to the bearing response without the LA14 component seen

in Figure D-2(c)), particularly in the x- direction induced by the strong coupled response

in the y- direction. Figure D-4 illustrates the response of a HDR bearing system to this

same LA13_14 table motion at a reduced span setting. As seen in the figure, the HDR

bearings exhibit a significant stiffening behavior, particularly in the first cycle excursions

to increasingly larger strain. The result of this behavior is increased force output.

However, reduced displacement response would be expected as a result. Figure D-5

illustrates the response of a system utilizing LR bearings to the LA13 component motion

at the same span setting as the previous test. As seen in the figure, the LR bearings exhibit

broad stable hysteretic response through numerous cycles (see Figure D-5(c)).

Comparisons between bearing characteristics (i.e., Qd, ku and kd) can be made by

106

examining the bearing hysteretic plots shown in these figures (see Table 3-1 and Figure 3-

1).

Figure D-6 through D-13 illustrate response for the FP, HDR, and LR bearing systems to

the LS17c_18c and the NF01_02 table motions. Comparisons between bearing behavior

are similar to the previous test results. The LS17c_18c motion produces relatively regular

harmonic response in the specimen, as noted in the displacement histories of Figure D-

6(a), D-7(a), and D-10(a). This soft-soil motion is rich in long period content (see Figure

3-10) near to the period characteristics of these test bearings, with harmonic response as

the expected result. Bi-directional coupling effects in the LS17c_18c motion are not

significant due to its strong orientation along its x-component. The near-fault NF01_02

motion has an impulsive fault-parallel (x-direction) and fault-normal (y-direction)

component which are inherently strongly coupled. This motion produces strongly coupled

behavior in the specimen as seen in the displacement orbits of Figure D-11 and D-12.

For scaled motion tests, representative results are presented in Figures D-14 through D-19.

These tests were performed utilizing FP bearings with type 1 PTFE composite. As seen in

the plots, the table motions at larger scale factors (i.e., lr) produce specimen response of

larger magnitude and longer period content (compare for example Figure D-14 to Figure

D-15). This difference in specimen behavior is to be expected, as the same motion at a

larger length scale factor produces typically larger spectral response (as seen in Figure 3-

11), particularly in the longer period range representative of the characteristic isolation

periods of the test bearings (see Table 3-1).

107

For near fault pulse tests, representative results are presented in Figures D-20 through D-

28. These tests were performed utilizing FP bearings with type 2 PTFE composite and LR

bearings. As seen in the plots, bearing response to these pure pulse motions are smooth

and cyclic. Bearing response is also strongly influence by coupling behavior induced by

the ground motion signal phasing of fault normal and parallel components. These results

show that fault parallel component pulses with durations of 2/3, 1, and 1.5 times the fault

normal pulse duration (Tp) produce decreasingly smaller coupled response, respectively.

In particular, this is noted in the decreasing bearing response along the direction of the

fault parallel component.

Figures D-29 & D-30 present results of sinusoidal bearing characterization tests for FP

bearings with type 1 and 2 PTFE composite sliders, respectively. Section 3.14 presents a

summary of these and all other FP slider characterization tests.

3.5 Experimental Results for Configuration 2 (Non-elevated, Eccentric Mass)




one or two summary figures are presented. The first figure presents three plots of the

global displacement time-history, global force time-history, and the x-direction bearing

hysteresis (for quadrant Q1). For systems with mass eccentricity, the second figure

presents a summary of torsional response including: a comparison of hysteretic response

and displacement time-histories for the East and West ends of the bridge specimen (in the

direction of motion input orthogonal to the direction of mass eccentricity), and a time-

108

history of deck rotation.

The Configuration 2 specimen was tested utilizing the x-direction component of the three

earthquake table motions. Tests were performed at two or more amplitudes (or span

settings) with mass eccentricities in the transverse direction of 0, 5, and 10 percent of the

overall bridge length. Tests were performed utilizing FP bearings (with type 2 PTFE

composite sliders) and LR bearings. Selected results are presented only for tests run at

their maximum amplitude. The figures illustrate the tendency for an increase in system

rotational response and bearing displacements at one end of the span as mass eccentricity

is increased (i.e., e/L increases from 0 to 0.05 to 0.10). This is made apparent by

comparing rotation histories and bearing hysteretic plots at the west and east ends of the

specimen for the same ground motion input with e/L = 0.05 and 0.10. For the case e/L = 0,

negligible system rotation was exhibited, as expected.

3.6 Experimental Results for Configuration 3 (Non-elevated, Eccentric Stiffness)

Table C-3 in Appendix C presents a complete log of shake-table tests performed on



two summary figures are presented. The first figure presents the x-direction global

displacement time-history, global force time-history, and bearing hysteresis (for quadrant

Q1). The second figure presents a summary of torsional response including: a comparison

of hysteretic response and displacement time-histories for the East and West ends of the

bridge specimen (in the direction of motion input orthogonal to the direction of stiffness

eccentricity), and a time-history of global deck rotation.

109

The Configuration 3 specimen was tested transversely utilizing the x-direction component

of the three earthquake table motions. Isolation system stiffness eccentricity was provided

by installing LR bearings and HDR bearings on the West and East ends of the bridge,

respectively. Selected results are presented only for tests run at their maximum amplitude.

Test results for this sequence are indicative of the stiffness eccentricity inherent in the

isolation system. Characteristic bearing hysteretic properties are compared in Table 3.1. It

is apparent the LR test bearings have higher initial stiffness (i.e., larger ku) and somewhat

higher characteristic strength (Qu). On the other hand, the HDR test bearings have larger

second-slope stiffness (kd) than the LR bearings at similar peak strains. The HDR bearings

also exhibit scragging effects resulting in increased force output in initial cycles. These

variations produce eccentricity in system strength, and first- and second-slope stiffness,

resulting in torsional behavior. The tendency in the test sequences is for the HDR bearings

(at one end of the span) to exhibit smaller displacement response than the LR bearings (on

the opposite end). This is presumably the result of the larger second-slope stiffness of the

HDR bearings, even though the LR bearings are initially stronger and stiffer. For the first

test cycles to large strain (see Figure D-61 and 62 for the LA13_14 test motion), HDR

bearings exhibited first cycle scragging effects resulting in force output larger than the LR

bearings. In later tests after scragging had occurred and insufficient time had been allotted

between tests to allow recovery in the HDR bearings, LR and HDR bearings exhibited

similar force output (despite the larger displacement response of the LR bearings).

110

3.7 Experimental Results for Configuration 4 (Elevated Single-span with Sym-metric Stiffness)




two summary figures are presented. The first figure presents the global displacement and

force orbits and x- & y-direction bearing hysteresis (for quadrant Q1). The second figure

presents a summary of local substructure response including plots of overall, bearing, and

pier hysteretic response in the x- & y-direction (at quadrant Q1).

The Configuration 4 specimen was tested utilizing the three selected earthquake time-

histories. Tests were performed at two or more amplitudes (or span settings) and bi-

directional motion pairs were run simultaneously and as separate x- & y-direction inputs.

Tests were performed utilizing FP bearings with type 4 PTFE composite. Selected results

are presented only for the bi-directional tests run at their maximum table span setting.

Similar to the Configuration 1 specimen test results (see Section 3.4), these results

indicate that response is influenced strongly by ground motion characteristics, including

frequency content and directional orientation. The LA13_14 motion has significantly

higher frequency content (see Section 3.2) and is strongly oriented along a 45 degree axis

to the x-y direction. This results in specimen response with significant coupling in the x-

and y-directions and underlying cycling of higher frequency (see Figure D-67 and D-68).

The LS17c_18c motion contains lower frequency components (see Section 3.2) and is

oriented strongly along the x-direction. This results in specimen response with strong

orientation along the x-direction with little coupled response and low frequency harmonic

111

cycling (see Figure D-69 and D-70). The NF01_02 motion is an impulsive near-fault event

with strong coupled x- and y-components (see Section 3.2). The vertical component

included in this ground motion history also contains significant amplitude accelerations.

This results in similar response to that of the LA13_14 motion, with less underlying high

frequency contribution, and added fluctuations in the bearing friction force component of

the hysteretic response induced by the vertical acceleration contributions (see Figure D-71

and D-72).

In addition, system attributes play an important contribution to the response as well.

Substructure pier response is essentially linear in these tests, with only slight pinching

near the origin, as seen in the hysteretic plots (see Figure D-68, D-70, and D-72 (e) and

(f)). This pinching behavior is discussed further in Section 3.13. On the other hand,

bearing response is essentially bilinear (see Figure D-67, D-69, and D-71 (c) and (d)), with

pinching of varying magnitude. This pinching response may be seen as the result of

several factors: bi-directional coupling in the bearing yield surface, vertical load

fluctuations caused by overturning effects and/or vertical acceleration input (see NF01_02

test results), or vertical load redistributions resulting from kinematic shortening of pier

assemblies. The latter phenomenon is discussed in further detail in subsequent evaluations

in Chapter 4. Finally, it is noted that total specimen displacement response (at the deck

level) is distributed between the isolation bearings and the pier substructure assemblies in

proportion to their flexibilities. Each of these components resist nearly equal force

transmission (varying slightly as the result of substructure mass contributions) with total

displacement response being the sum of the two component displacement contributions in

series.

112

3.8 Experimental Results for Configuration 5 (Elevated Single-span with Un-symmetric Stiffness)



are presented in Appendix D as Figures D-73 through D-98.

The Configuration 5 specimen was tested with and without bracing utilizing the three


settings) and bi-directional motion pairs were run simultaneously and as separate x- and y-

direction inputs. The specimen was also tested with uni-directional components of the

earthquake table motions at alternate length scale factors (i.e., lr), similar to the

Configuration 1 tests (see Section 3.4). The NF01_02 motion was run with and without its

vertical signal. Tests were performed utilizing FP bearings with type 3 and 4 PTFE

composites. Selected results are presented only for tests run at their maximum table span

setting.

For the selected earthquake history tests shown in Figure D-73 through D-92, two

summary figures are presented. The first figure presents the global displacement and force

orbits and x- and y-direction bearing hysteresis (for quadrant Q1). The second figure


pier hysteretic response in the x- and y-direction (at quadrant Q1). The characteristics of

specimen response to the different ground motion histories is similar to the Configuration

4 tests (see Section 3.7 above), with variations in ground motion frequency content and

directionality strongly influencing specimen behavior. For braced specimen tests,

substructure pier response is again nearly linear and notably stiffer than the unbraced

113

condition (compare Figure D-74 (e) and (f) to Figure D-76 (e) and (f)). For the unbraced

specimen, this configuration has larger pier stiffness in the x-direction compared to the y-

direction. This property can be readily observed from the response figures (see, for

example, Figure D-76 (e) and (f) and Section 3.1.4 and 3.1.5). Substructure stiffness

characterizations for these configurations are presented in Section 3.13 below. Foremost,

these results indicate that the effect of increasing substructure flexibility is an increase in

both global and substructure peak displacements. It appears, however, that bearing

displacement response may be somewhat larger, smaller, or relatively unaffected by this

variation (compare for example Figure D-74 to D-76). Further, FP bearing type 3 slider

composite is shown to have a lower friction value than the type 4 PTFE composite (see

Section 3.14 below), resulting in larger global and bearing displacement response when

these sliders are installed (compare Figure D-82 to D-84 (a) and (b)). Finally, it is noted

that bi-directional substructure stiffness is uncoupled in this specimen design. The effect

of this unequal x- and y-direction substructure stiffness (most notably in the unbraced

Configuration 5 specimen) is to skew the systems’ bi-directional yield surface by

producing unequal yield displacements in these directions. This effects the character of

global system hysteretic coupling, which can be seen by comparing Configuration 4 and

Configuration 5 hysteretic response for the same input motion (compare Figures D-72 and

D-86 (a) and (b)).

For uni-directional scaled motion tests in Figure D-93 through D-98, x-direction

displacement and force histories are presented in lieu of orbits, y-direction bearing

hysteresis at quadrant Q1 is omitted, and the plots showing local global, bearing, and pier

response distribution are omitted. The results of these tests are similar to the scaled motion

114

tests of Configuration 1 (see Section 3.4), with the table motions at larger length scale

factors (i.e., lr) producing specimen response of larger magnitude and longer period

content (compare Figure D-94 and D-97). This difference in behavior is similar to before,

as the isolated specimen will experience larger spectral response for the same motion at a

larger length scale (as seen in Figure 3-11).

3.9 Experimental Results for Configuration 6 (Elevated Single-span with Eccen-tric Stiffness)



are presented in Appendix D as Figures D-99 through D-122. For each of the selected

tests, four summary figures are presented. The first figure presents the global

displacement and force orbits and x- and y-direction bearing hysteresis (for quadrant Q1).

The second and third figures present summaries of local substructure response including

plots of overall, bearing, and pier hysteretic response in the x- and y-direction at quadrant

Q1 and Q4, respectively. The fourth figure presents a summary of torsional response

including: a comparison of hysteretic response and displacement time-histories for the

East and West ends of the bridge specimen (in the direction of motion input, orthogonal to

the direction of stiffness eccentricity), and a time-history of global deck rotation.

The Configuration 6 specimen was tested utilizing the three earthquake table motions.

Tests were performed at two or more amplitudes (or span settings) and bi-directional

motion pairs were run simultaneously and as separate x- & y-direction inputs. The

NF01_02 motion was run with and without its vertical signal. Tests were performed

utilizing FP bearings in two configurations. The first slider configuration utilized type 3

115

composite on all four bearings. The second slider configuration utilized type 4 and 3

composite on the East (unbraced) and West (braced) end of the bridge, respectively, in an

effort to counteract the torsional effects of the substructure stiffness eccentricity. Selected

results are presented only for bi-directional tests run at their maximum amplitude.

The torsional response induced by the specimen’s substructure stiffness eccentricity is

apparent from these test results, with global x-direction displacement (orthogonal to

stiffness eccentricity) larger on the unbraced East end relative to the braced West end (see

Figure D-102, D-106, D-110, D-114, D-118, and D-122). On the other hand, global

displacements in the y-direction are similar on the East and West end of the specimen with

displacement compatibility enforced by the longitudinally rigid deck frame (compare, for

example, plot (b) of Figure D-100 vs. D-101 and Figure D-116 vs. D-117). Bearing

displacements in the x-direction are not, however, systematically larger or smaller on the

unbraced end compared to the braced end of the specimen. This implies that shear

response is similar on either end of the specimen span. Torsional response is then mainly

due to kinematic rotation of the substructure about a vertical axis caused by the difference

in displacement between the flexible unbraced piers on the East end relative to the braced

West end piers. Bearing displacements in the y-direction, however, are always larger on

the braced West end of the specimen due to redistribution enforced by displacement

compatibility imposed by the longitudinally rigid deck frame (compare, for example, plots

(c) and (d) of Figure D-100 vs. D-101 and Figure D-116 vs. D-117).

It is also evident from these results that the alternate slider configuration, with lower

friction FP type 3 PTFE sliders installed on the braced end of the specimen and higher

116

friction type 4 sliders installed on the unbraced end, has a tendency to reduce rotational

response due to the underlying substructure stiffness eccentricity. This is illustrated by

comparing torsional response to the same input motion for the system utilizing the

uniform slider configuration and this alternate un-symmetric slider configuration

(compare Figure D-102 vs. D-106, Figure D-110 vs. D-114, and Figure D-118 vs. D-122).

3.10 Experimental Results for Configuration 7 (Elevated Double-span)




tests, three summary figures are presented. The first figure presents the global

displacement and force orbits and x- & y-direction bearing hysteresis (for quadrant Q1).


plots of overall, bearing, and pier hysteretic response in the x- & y-direction at quadrant

Q1 and Pier 5, respectively.





utilizing FP bearings in four configurations. The first two utilized type 4 and 5 PTFE

composite sliders, respectively, uniformly on all four bearings. The third system utilized

type 4 composite sliders on the East and West braced ends of the span and type 5

composite sliders on the flexible central piers (i.e., Piers 5 and 6) in an effort to reduce

117

shear transfer to these components. The final system utilized type 5 composite sliders on

the East and West braced ends and type 4 composite sliders on the central piers (Piers 5

and 6) in an effort to balance shear forces across all bent lines. Selected results are

presented only for bi-directional tests run at their maximum amplitude.

The effect of ground motion characteristics on specimen response in these tests is similar

to the Configuration 1, 4 and 5 test sequences. As discussed previously, input frequency

content and directionality playing a significant role in determining system dynamic

behavior. See previous discussions in Section 3.4 and 3.7.

The Configuration 7 specimen represents a simple two-span bridge overcrossing, with

relatively rigid end abutments and a flexible central pier bent. It is notable from these

results that this specimen exhibits similar global response as the Configuration 1 and

braced Configuration 5 test specimens when subjected to the same input motion (compare,

for example, Figure D-3, D-73, and D-132 for the LA13_14 input motion). This behavior

would be expected, as each of these configurations has a similar relationship of total mass

to global force-deformation behavior, with Configuration 7 having only the added

complexity of a flexible central pier bent modifying the hysteretic behavior of this

component.

The typical pattern of displacement and force distribution behavior in this specimen can

be seen by comparing hysteretic behavior at the end abutment location (i.e., Q1) to

response at the central pier bent (i.e., Q5) (see for example Figure D-124 and D-125 for

the LA13_14 input motion). The tendency is for bearings at the braced end of the

specimen to contribute nearly all of the total displacement demand at these locations.

118

Larger shear demand is also attracted to the braced abutment ends. The central pier

locations experience similar total displacement demands as end bents, due to displacement

compatibility enforced by the in-plane flexural rigidity of the deck frame. However, the

flexible piers at these locations provide a significant contribution to the total displacement

demand with bearings contributing the remainder. However, kinematic shortening of these

central pier assemblies (which occurs through rigid body rotation of these elements as the

tip of the pier is displaced) effects loss of axial force at these locations. This phenomenon

results in loss of FP bearing stiffness whose friction and pendulum stiffness components

are axial load dependant. This results in a reduction in shear force transmission at these

locations. This behavior is evident by noting the hysteretic pinching which occurs in

bearing response at these locations (see, for example, Figure D-125 and D-128 (c) and

(d)). The implications of this kinematic effect are discussed in further detail subsequently

in Chapter 4.

Bearing characterization tests show that the type 4 and 5 slider composites exhibited

similar friction values during these tests (see Section 3.14). Consequently, it is evident that

due to this similarity the effect of different slider configurations on system response

characteristics was minimal during these test sequences (compare, for example, Figure D-

123, D-126, D-129, and D-132).

3.11 Experimental Results for Configuration 8 (Elevated Double-span with Eccen-tric Stiffness)




119

tests, five summary figures are presented. The first figure presents the global displacement

and force orbits and x- & y-direction bearing hysteresis (for quadrant Q1). The second,

third and fourth figures present summaries of local substructure response including plots

of overall, bearing, and pier hysteretic response in the x- & y-direction at quadrant Q1,

Pier 5, and quadrant Q4, respectively. The fifth figure presents a summary of torsional

response including: a comparison of hysteretic response and displacement time-histories

for the East and West ends of the bridge specimen (in the direction of motion input,

orthogonal to the direction of stiffness eccentricity), and a time-history of global deck

rotation.





utilizing FP bearings with type 5 composite. Selected results are presented only for bi-

directional tests run at their maximum amplitude.

The pattern of torsional response induced by the specimen’s substructure stiffness

eccentricity is similar to the Configuration 6 test sequences. Global x-direction

displacement (orthogonal to stiffness eccentricity) increases along the span moving from

the braced West end to the unbraced central piers to the unbraced East end (see, for

example, Figure D-160, D-161, D-162 and D-163 for the LA13_14 input motion). As

before, global displacements in the y-direction are similar on the East and West end of the

specimen with displacement compatibility enforced by the longitudinally rigid deck

120

frame. Bearing displacements in the x-direction are similar implying similar shear

response along the specimen span. Torsional response is then mainly due to kinematic

rotation of the substructure about a vertical axis caused by the difference in displacement

between the piers along the span. Bearing displacements in the y-direction, however, are

always larger on the braced end of the specimen due to pier stiffness at this location and

displacement compatibility imposed by the longitudinally rigid deck frame.

3.12 Experimental Results for Configuration 9 (Elevated Double-span with Yield-ing Piers 5 and 6)













utilizing FP bearings. Type 5 PTFE composite sliders were utilized on the East and West

braced ends of the specimen while type 4 sliders were utilized on the central piers (i.e.,

Piers 5 and 6). The vertical load was also balanced such that approximately 5/8ths was

121

supported by Piers 5 and 6 (consistent with the loading of a continuous girder span) in an

effort to attract shear forces to this location.

The Configuration 9 specimen configuration is essentially identical to Configuration 7,

with the addition of weaker A36 leaf springs installed in the central piers. Several test

sequences run at increasing span settings were performed on this specimen with limited

yielding occurring in central piers (see, for example, pier hysteresis in Figure D-179 (e)

and (f)) Consequently, pier bracing at the end abutments was removed in an effort to

balance shear behavior across the specimen increasing demands at the yielding central

piers. In this unbraced configuration, span settings were again increased to maximize

ductility demand at the central piers (see pier hysteresis in Figure D-181 and D-184 (e)

and (f)). Peak ductility demand of approximately 1.5-2 were achieved in the final tests

sequence with no apparent reduction in the isolated performance of the system (see Figure

D-182, D-183, and D-184).

3.13 Experimental Results for Configuration 10 (Substructure Static Pullback Tests)

Table C-10 in Appendix C presents a complete log of static pull-back tests performed on

Configuration 10 of the isolated bridge specimen. Selected experimental results from

these tests are presented in Appendix D as Figure D-185. Braced frame stiffness

characterization was computed from dynamic test data (see, for example, Figure D-178 (e)

and (f)). Table 3-7 below tabulates substructure pier assembly properties computed from

these test results.

122

Substructure properties in Table 3-7 determined from these characterization tests can be

compared to the previous computed design values shown in Table 3-2. As seen, a number

of discrepancies between the design and test values are apparent. First of all, stiffness

values are generally lower than computed design values. Actual braced pier stiffness is

slightly less than computed (i.e., 26.7 kips/inch computed vs. 22-24 kips/inch tested). This

reduction is apparently due to friction slip of bolted end connections. In addition, actual

stiffnesses of short and long plate pier assemblies are lower than computed (i.e., 4 kips/

inch computed vs. 2.4-3 kips/inch tested for the short plate assembly and 1.8 kips/inch

computed vs. 1.6-1.7 kips/inch tested for the long plate assembly). These stiffness

reductions are likely due to gapping in the mechanical assemblies which occurred at

movements near the displacement origin. Further, assumed slider elevations were slightly

lower than as-built elevations in design computations. This led to a higher computation of

effective lateral stiffness in design due to the displacement transformation at the lower

assumed slider height. Finally, it is seen that strength values for short and long plate pier

assemblies were higher than computed (i.e, 2.9 kips design value vs. 4 kips tested value).

This effect was certainly the result of material strength above the assumed nominal value

(i.e., 36 ksi for A36 material).

3.14 Experimental characterization data for FP slider composites

A number of tests were performed during the shake table simulation sequences to evaluate

the friction coefficients of the five FP bearing slider composites provided by the

manufacturer. A log of these bearing characterization tests are shown in Table C-1, C-4,

C-5, and C-7 of Appendix C.

123

The shear response, V(u), of a spherical sliding FP bearing responding uni-directionally

can be represented by the following,

(14)

where µ is the friction coefficient of the slider-dish interface, u the uni-directional

displacement, the velocity, N the normal force (possibly varying in time), and R the

bearing radius [Constantinou et al. 1998]. It has been shown that the friction coefficient of

a PTFE slider moving across a stainless steel surface increases with velocity up to a

threshold value [Constantinou et al. 1990] (see Figure 3-14 below). Bearing

characterization tests were performed by applying test signals to the bridge specimen

which produced essentially harmonic bearing response through a range of velocities (see

Figure 3-13 and Table 3-6). Bearing shear response data from these tests were processed

Table 3-7 Substructure pier assembly properties computed from characterization data

Configuration1 Material Ksub (kips/in) α2 yield point 3

(kips)Fy 4

(kips)Tsub

5(sec)

Braced x- 22 0.27

Braced y- 24 0.26

Short Plate A514 2.4-2.6 0.83-0.80

Long Plate A514 1.6-1.7 0.99-1.02

Short Plate A36 2.4-3 .04-.167 4 4.75-5 0.83-0.74

Long Plate A36 1.6-1.7 .23-.25 4 4.6 0.99-1.02

1. see Section 3.1.4

2. strain hardening ratio, see Figure D-185

3. first point of yield onset, see Figure D-185

4. vertex of bilinear idealization, see Figure D-185

5. based upon tributary mass; see Chapter 2, Equation 3

V u( ) µN u·( )sgn NR----u+=

u·

124

to remove the stiffness contribution (i.e., N/R) and then divided by the time varying

normal force to determine the friction coefficient of each slider composite. Mean friction

coefficient results for each of the slider composite types plotted as a function of velocity

are shown in Figure 3-15. Error bars are included in the plots indicating one standard

deviation statistical scatter.

As seen from these results, friction coefficients ranged from lowest to highest for the type

1, 3, 4, 5, and 2 slider composites, respectively. These composite types (in their virgin

condition) exhibited peak average friction coefficient values of approximately 0.05,

0.0575, 0.08, 0.095, and 0.103, respectively. The type 1 slider composite exhibited a

significant increase in friction coefficient after 57 test signals (see Figure 3-15(a)). The

type 4 slider composite exhibited a similar increase in friction response, after 231 tests

were performed (see Figure 3-15(d)). Slider composite type 5, on the other hand, exhibited

a slight reduction in average friction response after 48 tests (see Figure 3-15(e)).

Figure 3-15(f) illustrates the hysteretic friction response of bearings using the type 5 slider

composite subjected to the LS17c table motion time-history. It is apparent from this figure

that higher friction response (approaching 13 percent) is exhibited in the initial

displacement cycle of the bearing. This behavior may indicate an initial “stick”

phenomenon in the slider interface not noted in the previous characterization tests which

were not processed to near zero velocity. As seen in the remaining response, however,

friction of near 10 percent is exhibited during peak displacement cycles. This is consistent

with the bearing characterization data (see Figure 3-15(e) for values of the peak average

friction coefficient plus one standard deviation). As displacement cycles subside in the

125

remaining hysteretic response, friction approaches 8%, consistent with the lower velocity

response for this composite (see Figure 3-15(e)). This is consistent with reported behavior

due to rate effects (see Figure 3-14) [Constantinou et al. 1990].

Figure 3-14 PTFE slider composite behavior as a function of velocity

velocity

fric

tion

coef

ficie

nt

126

a. Type 1 b. Type 2

c. Type 3 d. Type 4

e. Type 5 f. Type 5 friction hysteresis: LS17c

Figure 3-15 Characterization of FP bearing slider composites

0 1 2 3 40.03

0.04

0.05

0.06

0.07

0.08

0.09

velocity (in/sec)

fric

tion

coef

ficie

nt

virgin

after 57 tests

0 1 2 30.06

0.07

0.08

0.09

0.1

0.11

0.12

velocity (in/sec)

fric

tion

coef

ficie

nt

virgin

0 1 2 3 40.045

0.05

0.055

0.06

0.065

0.07

velocity (in/sec)

fric

tion

co

effic

ien

t

virgin

0 1 2 3 40.04

0.06

0.08

0.1

0.12

velocity (in/sec)

fric

tion

coef

ficie

nt

virgin

after 231tests

0 1 2 3 4 5

0.08

0.09

0.1

0.11

velocity (in/sec)

fric

tion

coef

ficie

nt

virgin

after 48tests

−3−2.5−2−1.5−1−0.5 0 0.5 1 1.5 2 2.5 3−0.15

−0.1

−0.05

0

0.05

0.1

0.15

displacement (in)

fric

tion

co

effic

ien

t

67

3 Testing of a Bridge Deck Model

3.1 Introduction

The experimental studies in this research were developed to satisfy the program objectives

outlined in Section 1.3.1. Specifically, these studies fulfill program Task 6 in Section

1.3.2, to construct a scale model of a bridge with flexible piers for earthquake testing.

These tests utilized the newly upgraded 3-dimensional capabilities of the earthquake

simulator at the Pacific Earthquake Engineering Research (PEER) center and provided a

first-time opportunity to study the bi-directional and 3-dimensional response of isolated

bridge systems.

The preliminary analytical evaluations of Chapter 2 illustrate the basic sensitivity of the

response of isolated bridge overcrossings to various structural and isolator characteristics.

Based on these evaluations. the specimen tested were developed to confirm these

sensitivities, including the effect of mass and stiffness eccentricity, pier flexibility, pier

mass and strength, and varying isolator properties, on system response.

These simulation tests also subjected various isolated bridge overcrossing configurations

to multiple signals of one to three components of earthquake ground motion. These

earthquake signals were representative of significant far-field, near-fault, and soft-soil

seismic events. In this respect, these studies provide invaluable proof-of-concept

68

verification for the effectiveness and robustness of seismic isolation for simple bridge

types.

3.1.1 Similitude Requirements

Similitude requirements for dynamic structural models have been developed in previous

studies [e.g., Krawinkler, Moncarz, 1982]. These similitude constraints are defined by the

chosen length scale factor, 1r, prescribed for relating model size to the subject prototype

(full-scale) structure (i.e., lr = lm/lp, where lm and lp are the model and prototype length

scale, respectively). With gravity a constant and masses reduced proportionally by the

length factor, ground motion accelerations are held constant in dynamic testing to preserve

force scaling (where F = ma). Similitude requires that prototype and model periods be

related by the square root of the length scale factor (i.e., the time scale factor, tr = 1r1/2),

with frequencies related by its inverse (i.e, the frequency scale factor, ωr = lr-1/2). Ground

motion acceleration time-history records must then be “compressed” by multiplying their

time increment by the time scale factor to preserve similitude. This produces time-history

records for shake table simulations of equal acceleration output (as the original recorded

motion), but of higher frequency. With these constraints applied specimen simulations are

performed, with model and prototype system displacements (i.e., dm and dp, respectively)

related directly by the length scale factor (i.e., the displacement scale factor, dr = dm/dp =

1r) [Krawinkler, Moncarz, 1982].

In Chapter 2, isolated bridge behavior was shown in several cases to vary with system

properties corresponding to “short” and “long” spectral period ranges. The typical

response of these systems was shown in several cases to have unique characteristics in

69

each of these period ranges. To model this behavior experimentally, similitude requires

that the model period, Tm, be related to the prototype system period, Tp, by the time factor

(i.e., tr = 1r1/2) such that

(12)

With this requirement, model stiffnesses (Km) can be selected to produce required model

periods (i.e., Tm), given a model mass (Mm), through the definition of system period as

(13)

A length scale factor of 1/4 (i.e., lr = 1/4) was chosen for model scaling for these studies.

This ratio was selected based upon the load and geometric constraints of the earthquake

simulator which were reached for bridge model weight and length quantities represented

by a typical single-span bridge overcrossing reduced to this size. Ground motion scaling

was performed separately, with length scale factors chosen between 1/2 to 1/5 (i.e., lr

between 1/2 and 1/5) and related time scale factors between and (i.e., tr = 1/ and

1/ , see Equation 12), to produce ground motion simulation records of peak velocity and

displacement within the capacity of the shake table simulator (see Section 3.2 below).

The resulting relationships between model and prototype response produced by tests using

these separate model and ground motion scaling factors may be deduced by applying the

previous similitude relations. For example, for a given model with length scale factor of

trTmTp------- lr

1 2⁄==

Km 4π2MmTm--------=

2 5 2

5

70

lrm, a full-scale prototype system with a period of Tp = Tmlrm-1/2 is represented (see

Equation 12). Applying a different length scale factor of lrg to the ground motion implies

that this same model represents a prototype system with a period of Tp’ = Tmlrg-1/2

subjected to this scaled ground motion record (Equation 12). Combining these provides

Tp’ = Tp(lrm/lrg)1/2. Thus if the ground motion scale is smaller than the model scale (e.g.,

lrg = 1/5 and lrm= 1/4), testing with this ground motion represents the response of a

prototype system with a period, Tp’, larger than the full scale target prototype, Tp. On the

other hand, if the ground motion scale is larger than the model scale (e.g., lrg = 1/2 and

lrm= 1/4), the testing represents the response of a prototype system with a period, Tp’,

smaller than the full scale target prototype, Tp. In this way, the model (with fixed scale,

e.g., lrm = 1/4) may represent systems in different spectral regions of the ground motion

record by varying the chosen scale factor (i.e., lrg) applied to the motion in the given test

sequence. In either case, displacements to the full-scale motion are related to model

displacements in the test sequence by the ground motion scale factor, lrg, as before.

3.1.2 Design criteria

In order to illustrate the effect of basic system variations on the response of the isolated

bridge specimen, component properties were established to accommodate the same

substructure and isolation system variations as those evaluated in Chapter 2. These

properties were established to model systems over the broad spectral range where

preliminary analyses indicated significant variations in response.

71

Bridge size and mass were chosen to produce the largest similitude length scale within the

limits of the shake table physical geometry and vertical load capacity. This implied a ¼-

scale model (i.e., length scale factor, lr = 1/4) with maximum span of 20 feet and a

corresponding bridge deck mass of 65 kips.

Target isolation system properties were determined for the given bridge deck mass. It was

established that model isolation bearings with characteristic periods of Tiso = 1.3 to 1.75

seconds were desired (see Equation 2). These would represent full-scale (prototype)

isolated systems with a rigid-based isolation periods ranging from Tiso = 2.6 to 3.5 seconds

at lr = 1/4. Isolator strengths were selected to be in the typical range of Cyiso = 0.04 to 0.12

(see Equation 1). Several bearing types were to be represented to provide hysteretic

properties covering these ranges, including Lead-rubber (LR), High-damping rubber

(HDR), and Friction Pendulum (FP) bearing types.

Test specimen non-isolated periods of approximately Tsub = 0.025 to 1 seconds were

desired (see Equation 3). These represent prototype (full-scale) non-isolated bridges with

periods ranging from Tsub = 0.05 to 2 seconds at a length scale factor of lr = 1/4 which

covers the range of substructure systems evaluated in Chapter 2. The flexibility of

specimen substructure components was established to target substructure stiffness (Ksub)

to isolator second-slope stiffness (kd) ratios in the range of 2:1 to 7:1. As the AASHTO

Guide Spec presumes some measure of substructure yielding to occur (see Section 2.2), it

was also considered important to study its effect on the response of isolated bridge

systems. Therefore, weaker specimen substructure components designed to yield near

peak response were desired as part of the test specimen arrangement.

72

A versatile specimen which could accommodate these variations in pier flexibility and

isolation bearing properties was required for these studies. In addition, the model had to

accommodate varying pier mass and strength and eccentric arrangements of superstructure

mass and substructure stiffness. The design also needed to be easily alterable during

testing to accommodate these system changes efficiently.

3.1.3 Design Development

In keeping with these requirements, a ¼-scale bridge specimen design was developed with

preliminary consideration given to several alternate systems. Steel columns were

preferred because of the desire to perform many tests on the specimen without damage to

the total substructure. A simple design utilizing standard pipe column sections with

optional cross bracing (added for rigid or eccentric configurations) was first considered.

Standard pipe material (i.e., nominal yield strength, Fy = 46 ksi) proved inadequate to

resist combined effects of axial and flexural loads at pipe lengths required to achieve

target substructure flexibility. A variety of other ideas were then considered, including an

alternative utilizing steel wide-flange, L-frames mounted on 3-dimensional clevis base

pins. For this alternative, prestressed coil spring assemblies were designed to provide

variable rotational base flexibility. These assemblies allowed variations in the lateral

frame stiffness to be achieved by replacing springs of pre-engineered stiffness in the base

connection assembly. Connection detailing, necessary to ensure smooth movement of the

assemblies through the anticipated displacement ranges, proved complicated, however,

causing this alternative to be abandoned in lieu of the simpler final specimen design. Final

specimen design was taken as a combination of these two alternatives, as described in the

next section.

73

3.1.4 Final Specimen Design

Complete construction plans of the test specimen, including as-built photographs, are

included in Appendix A for reference. The final specimen substructure design utilized

steel tube frames mounted on clevis base pins. The clevis pin mounts rotate about two-

orthogonal axes and provide adequate transverse and longitudinal rotation capacity at the

base (with restraint about the vertical axis). Rotational stiffness was provided by thin

tapered plates extending in each orthogonal direction from the base of the steel tube

frames. The wide end of these plates attach to the tube frames near the clevis pins while

their narrow tapered ends attach to a shear pin connection which provides vertical support

only. Tube columns support load cells and isolation bearings, which in turn support a rigid

steel deck frame spanning between pier assemblies. The tube column frames support

vertical gravity load (imposed by the 65 kips of concrete blocks mounted on the bridge

deck frame as mass) and provide lateral stiffness supporting shear loads transmitted from

the bridge deck through the isolation bearings during seismic excitation.

Two tapered plate designs were prepared for the elevated bridge specimen. These two

designs allowed substructure lateral stiffness to be altered by changing the rotational

flexibility at the base of the piers (see Appendix A). Short and long tapered plates

provided the specimen with non-isolated periods of approximately Tsub = 0.67 and 1

seconds, respectively, Adding cross-bracing between the columns provided the specimen

with a non-isolated period of approximately Tsub = 0.25 seconds. Two sets of each size

tapered plates were fabricated from A514 and A36 steel, respectively. The stronger A514

material provided essentially elastic stress levels, while the weaker A36 material allowed

moderately ductile substructure behavior, during the test sequences.

74

Three separate sets of test bearings were designed and fabricated for use in the bridge deck

model tests: 1) high-damping rubber (HDR) bearings supplied by Bridgestone, Inc.; 2)

lead-rubber (LR) bearings supplied by DIS, Inc.; and 3) spherical sliding, or Friction-

pendulum (FP) bearings supplied by Earthquake Protection Systems, Incorporated.

Design details of each are included in Appendix A for reference.

The basic force diplacement properties of these three bearing types (see Figure 3.1) were

intentionally selected to be different in order to cover the ranges of isolation bearing

properties normally seen in practice and to examine behavior in different ranges of

behavior identified in the preliminary evaluations (see Chapter 2). One should not use

these tests to compare the effectiveness of different isolator types, as different designs

could also have resulted in isolators with very similar hysteretic properties. Results from

these tests should be used only for calibration of models, validation of identified trends

and behavior concepts (i.e., how system behavior is effected by changes in the

fundamental hysteretic properties Qd, ku and kd), and the efficacy of isolation in general.

Bearing hysteretic properties established from characterization test data are presented in

Table 3-1 as a reference. Figure 3-1 shows a uni-directional plot comparing the hysteretic

response of the HDR, LR, and FP test bearings illustrating the hysteretic differences in

strength and first- and second-slope stiffness characteristics of these bearing designs.

Characteristic target (rigid-based) isolation periods (i.e., Tiso, see Equation 2) are also

shown in Table 3-1 for each bearing type.

As illustrated in the plans (see Appendix A), this final specimen design is quite versatile.

Lateral substructure stiffness can be distributed uniformly at each span or

75

unsymmetrically (with increasing flexibility along the bridge length) to model bridges

spanning grade changes. Substructure flexibility can also be varied in the longitudinal and

transverse directions to model the cantilever versus frame behavior of a multi-column

bent. For two-span (six-pier) configurations, end piers can be braced longitudinally and

transversely and the central piers left relatively flexible to model bridge overcrossings.

Table 3-1 Summary of bearing mechanical properties from characterization tests

Bearing Type Qd (/W) ku(kips/in) kd (kips/in) Tiso4 (sec)

HDR1 0.015-0.05 4.5-3.5 1.75-1 1.29

LR2 0.05-0.09 15-5 1.6-0.9 1.36

FP3 µ = 0.07 - 0.10 INF R = 30 in.W/R= 0.54 kip/in 1.75

1. Property range over 50-250% shear strain with Tiso at γ = 250%2. Property range over 50-200% shear strain with Tiso at γ = 200%3. FP hysteretic properties based on friction coefficient, µ, and dish radius, R4. Tiso computed for a deck weight of 65 kips applied uniformly to four bearings

Figure 3-1 Comparison of test bearing hysteretic properties

0

0.05

0.1

0.15

0.2

0.25

0 0.5 1 1.5 2

d (in)

F/W

HDRLRFP

76

A column detail also allows additional substructure mass of up to 1.5 kips to be installed

near the 2/3rd height of each column frame. This represents substructure mass proportions

of up to γ = Σ mpier/Mdeck = 10 percent.

For some configurations, the isolators are mounted directly on the shake table. Separate

transition hardware allows FP, LR and HDR isolation bearings to be installed. Bridge deck

mass can also be arranged symmetrically or eccentrically. The mechanical properties of

the bridge specimen substructure are presented in Table 3-2.

Table 3-2 Summary of computed specimen substructure design properties

Frame Configuration

Braced1 Short Plate2 Long Plate3

Tsub 0.25 sec 0.67 sec 1.0 secKsub 26.6 kips/in 4 kips/in 1.8 kips/in

Fy N.A. 8 kips (A514)2.9 kips (A36)

8 kips (A514)2.9 kips (A36)

Cy = Fy/W N.A. 0.5 (A514)0.18 (A36)

0.5 (A514)0.18 (A36)

dy N.A. 1.9 in (A514)0.68 in (A36)

4.0 in (A514)1.44 in (A36)

Ksub/kd (LR) 27:1 4:1 2:1Ksub/kd (FP) 47:1 7:1 3:1

1. Braced pier assembly- see Figure A-1, A-2, & A-7 details, Appendix A2. Pier assembly with short plate - see Figure A-1, A-2, & A-10, Appendix A3. Pier assembly with long plate - see Figure A-1, A-2, & A-10, Appendix ATsub - see Chapter 2, Equation 3Ksub - lateral stiffness at top of pier assemblyFy, Cy - yield strength and yield strength coefficient of pier assemblydy - yield displacement at top of pier assemblyKsub/kd - substructure stiffness to isolator second-slope stiffness ratio

77

3.1.5 Test Set-Up and Protocol

Utilizing the bridge specimen substructure components and isolation test bearings

provided, ten bridge configurations were developed to study various aspects of the

response of isolated bridge overcrossings through earthquake simulation studies. This test

program was developed to examine the global and local response trends of simple and

more complex isolated bridge systems subjected to a variety of earthquake input,

fullfilling objective 2 and 3 of the research program (as outlined in Section 1.3.1). Specific

issues to be examined in the study were outlined in project Task 6 (see Section 1.3.2) and

discussed further in the preliminary evaluations of Chapter 2. These issues include the

effect on isolated bridge response of: (1) the characteristics of the earthquake input

(including fault proximity, site soil effects, and the effect of uni-directional vs. multi-

directional input); (2) isolation system characteristics; (3) substructure flexibility; (4)

substructure mass; (5) higher modes; (6) system eccentricity; and (7) substructure

yielding. The ten bridge specimen configurations developed for these studies are

described in the sections below and schematically illustrated in Figure 3-2 through 3-6.

Configuration 1

This configuration was the standard non-elevated symmetric mass configuration of the

bridge deck model (see Figure 3-2). Table motion in this case represents motion at “rigid”

abutments. In this configuration, load cells were mounted directly on the earthquake

simulator. Isolation bearings were mounted to the top of load cells and connected to the

underside of the steel bridge deck. Concrete blocks were prestressed to the deck frame,

providing 65 kips of weight. Instrumentation for this configuration included: (1) table

78

instrumentation measuring table accelerations and displacements; (2) five degree-of-

freedom load cells measuring x- & y- shear and moment, as well as axial forces, under

each isolation bearing (at quadrant Q1, Q2, Q3 and Q4); (3) linear potentiometers (LP)

measuring bearing, deck, and table displacements; (4) DCDT’s measuring potential

shifting of the concrete mass blocks ; and (5) accelerometers mounted to the bridge deck,

concrete mass, and locally above isolation bearings. Data acquisition in this configuration

is illustrated in Figure 3-7 and detailed descriptions of each instrumentation device are

outlined in Table 3-3.


overcrossings. Variations in isolator characteristic properties (i.e., first- and second- slope

stiffness and strength) were examined in this configuration by replacing bearing types.

The effect of earthquake motion characteristics, including bi-directional and vertical input,

fault-proximity and soil effects, were studied in this and all model configurations. The

effect of idealized near-fault pulse motions were studied solely in this basic configuration

(see Section 3.2 for information regarding earthquake test motions).

Configuration 2

This configuration was similar to Configuration 1, but incorporated an eccentric mass

layout for the bridge deck (see Figure 3-2). This configuration was erected similar to

Configuration 1 with concrete blocks offset to produce superstructure mass eccentricities.

Instrumentation for this configuration was the same as that for Configuration 1 (see Figure

3-7 and Table 3-3).

79


overcrossings. Deck mass was systematically offset to produce superstructure mass

eccentricities of 5 and 10 percent of the longitudinal bridge span. Torsional response of

different isolation systems were also examined in this configuration by replacing bearing

types.

Configuration 3

This configuration was again similar to Configuration 1, but incorporated an unsymmetric

distribution of isolation bearings (see Figure 3-3). This configuration was erected similar

to Configuration 1 with two different isolation bearing types (HDR and LR bearings)

mounted at either end of the bridge span. Instrumentation for this configuration was the

same as that for Configuration 1 (see Figure 3-7 and Table 3-3).

Configuration 3 was developed to study the torsional response of isolated bridge

overcrossings produced by bearing stiffness eccentricity. The installation of HDR and LR

bearings at either end of the deck produced eccentricity in first- and second- slope stiffness

and bearing strength properties. This study was not performed to suggest that different

types of isolators could be installed in a bridge at opposite abutments, but to assess the

effect of these types of variations along a bridge span. The effect of earthquake input

characteristics on this stiffness induced torsional response were also studied in this

configuration.

80

Configuration 4

This was the standard four-pier elevated configuration of the bridge deck model, an

elevated single-span configuration with symmetric substructure stiffness (see Figure 3-3).

In this configuration, substructure pier assemblies were mounted directly to the

earthquake simulator with load cells mounted on top of pier supports. FP isolation

bearings were used exclusively in this configuration attached to the top of load cells and

connected to the underside of the steel bridge deck. Concrete blocks were prestressed to

the deck frame, providing 65 kips of weight. In addition to the same instrumentation

utilized for Configuration 1 through 3 (see Figure 3-7 and Table 3-3), the following

instrumentation was added in this configuration: (1) additional linear potentiometers (LP)

measuring top of pier displacements; (2) additional accelerometers mounted to top of

substructure piers; and (3) linear strain gauges mounted on tapered leaf springs at base of

piers to monitor potential yield conditions. Data acquisition hardware in this configuration

is illustrated in Figure 3-7, 3-8, and 3-9. Detailed descriptions of each instrumentation

device are outlined in Table 3-3 and 3-4.


overcrossings and viaducts. Variations in isolator characteristic strength were examined in

this configuration by replacing FP bearing slider types. The effect of variations in

substructure flexibility were examined by adding or removing bracing or using different

rotational leaf springs at the base of the pier assemblies.

81

Configuration 5

This was an elevated bridge configuration with un-symmetric substructure stiffness,

erected similarly to Configuration 4, with short and long leaf springs inserted at the base

of pier assemblies in the x- and y- directions, respectively (see Figure 3-4).

Instrumentation for this configuration was similar to Configuration 4 as illustrated in

Figure 3-7, 3-8, and 3-9 and outlined in Table 3-3 and 3-4.


overcrossings, with the effect of variations in substructure flexibility in the x- and y-

directions. Variations in isolator characteristic strength were examined in this

configuration by replacing FP bearing slider types as before. The effect of substructure

mass (where γ = Σ mpier/Mdeck) was also studied in this configuration by varying the

number of attached lead weight packets mounted to substructure pier assemblies.

Configuration 6

This configuration is similar to Configuration 5 with additional cross-bracing added to

piers at one end of the specimen (see Figure 3-4). Instrumentation for this configuration

was similar to Configuration 4 and 5 as illustrated in Figure 3-7, 3-8, and 3-9 and outlined

in Table 3-3 and 3-4.


overcrossings having unsymetric substructure conditions. Variation of substructure

flexibility along the span produced by added cross-bracing created the effect of

eccentricity in substructure stiffness in the configuration. Variations in isolator

82

characteristic strength could be examined in this configuration by replacing FP bearing

slider types symmetrically (i.e., replacing all slider types at once) or unsymmetrically (i.e.,

placing weaker FP bearing sliders above braced piers to counteract torsional response).

Configuration 7

This was the standard elevated double-span configuration of the bridge deck model (see

Figure 3-5). In this configuration, six substructure pier assemblies were mounted directly

to the earthquake simulator. Cross-bracing was added to piers in the longitude and

transverse directions at each end of the bridge specimen to simulate stiff abutment

conditions. Center piers were configured with short and long leaf springs inserted at their

base in the x- and y- directions, respectively, simulating a flexible central pier bent. Load

cells were mounted on top of pier supports with FP isolation bearings attached to the top

of load cells and connected to the underside of the steel bridge deck. Concrete blocks and

additional lead packets were prestressed to the deck frame providing 99 kips of weight.

This additional mass was installed to provide the system with an overall yield strength

coefficient (i.e., Cyiso) theoretically equivalent to the four pier specimen (assuming a

constant FP slider µ value). In addition, this larger mass produced nearly equivalent period

characteristics for the six-pier configuration with out cross-bracing as the four pier un-

braced configurations. The instrumentation for this configuration was similar to

Configuration 4 through 6, with the following additions: (1) five degree-of-freedom load

cells under each isolation bearing at pier 5 and 6; (2) linear potentiometers (LP) measuring

top of pier 5 and 6 displacements; (3) accelerometers mounted to locally above isolation

bearings and at top of substructure piers at pier 5 and 6; and (4) linear strain gauges

83

mounted on tapered leaf springs at base of pier 5 and 6. Data acquisition hardware is

illustrated in Figure 3-7, 3-8, and 3-9 and outlined in Table 3-3 and 3-4, as before.


bridge overcrossings. Variations in isolator characteristic strength were examined in this

configuration by replacing FP bearing slider types. These strength variations were

implemented symmetrically (i.e., by replacing all sliders at once) or unsymmetrically (by

placing weaker FP sliders selectively above braced piers or central piers to effect force

distribution).

Configuration 8

This configuration was similar to Configuration 7 with cross-bracing added to piers at

only one end of the specimen and long tapered leaf spring plates installed at the base of

piers in the x- direction at the opposite end (see Figure 3-5). Instrumentation for this

configuration was similar to Configuration 7 as illustrated in Figure 3-7, 3-8, and 3-9 and



overcrossings. Increasing substructure transverse flexibility along the span produced

eccentricity in substructure stiffness. The effect of varying isolator characteristic strength

along the span was examined by replacing FP slider types symmetrically (i.e., replacing

all slider types at once) or unsymmetrically (i.e., placing weaker FP bearing sliders above

braced piers to counteract torsional response).

84

Configuration 9

This configuration was similar to Configuration 7 with weaker (A36) tapered leaf spring

plates inserted at the base of central piers (see Figure 3-6). Instrumentation for this

configuration is similar to Configuration 7 as illustrated in Figure 3-7, 3-8, and 3-8 and



bridge overcrossings with yielding substructure components. A36 leaf spring plates in

central piers were designed to allow ductile response in the lateral force-displacement

characteristics of the central pier assemblies.

Configuration 10

This was a free-standing configuration of substructure pier sub-assemblies (see Figure 3-

6). This configuration utilized the same substructure assemblages as Configuration 9 with

bridge mass and deck, isolation bearings, and load cells removed. Instrumentation for this

configuration consisted of: (1) a linear load cell attached in series with a come-along

winch between tops of pier sub-assemblies to produce an increasing static force, and (2)

linear potentiometers (LP) measuring the displacements at the top of the piers (see Figure

3-6). This configuration was developed to characterize the lateral force-deformation

characteristics of the strong and weak substructure pier sub-assemblies (i.e., those

utilizing A514 and A36 tapered leaf spring base plates, respectively). The piers were

characterized in both their stiff (i.e., short plate) and flexible (i.e., long plate)

configurations (see Table 3-2).

85

Specimen as-built documentation

As a reference, photographs of bridge specimen erection sequences, pier frame

assemblies, and component details are included along with construction plans and

fabrication details in Appendix A.

3.2 Earthquake Histories for Testing

An evaluation of the effect of ground motion input characteristics on the response of the

various configurations of the bridge deck specimen was an essential component in these

simulation studies. The features of seismic demand include: (1) fault proximity (i.e., far-

vs. near- field events); (2) directivity (i.e., fault-normal vs. fault-parallel and forward- vs.

backward-azimuth motions); (3) site specific soil conditions (i.e., rock vs. soil sites); (4)

bi-directional effects; and (5) vertical components of motion. A group of shake table

motions was tailored to consider each of these variations explicitly. In addition, sinusoidal

forcing motions were developed to characterize the friction coefficient of the various FP

bearing slider types provided by the manufacturer. The characteristics of these motions are

described in detail below.

Three representative earthquake time-histories were selected for shake table testing. Basis

recorded time-histories were chosen from the database developed in Phase 2 of the

FEMA/SAC project [SAC, 1997]. The bi-directional ground motion pairs LA13/LA14,

NF01/NF02, and LS17c/LS18c, as listed in Table 2-2, were selected from this database to

represent far-field, near-fault, and soft-soil ground motion types, respectively. These basic

motions were filtered and scaled to produce shake table input motions conforming to the

acceleration, velocity, and displacement limits of the shake table apparatus. The resulting

86

Configuration 1: Non-elevated Symmetric Mass

Configuration 2: Non-elevated Eccentric Mass


x

y

(4) FP bearingsor(4) LR bearingsor (4) HDR bearings

C.M.


x

y

(4) FPS bearings

C.M.

ey

or(4) LR bearings


87

Configuration 3: Non-elevated Eccentric Stiffness

Configuration 4: Elevated Single-span with Symmetric Stiffness


x

y

(2) HDR bearings

(2) LR bearings


C.M.

x

y

(4) FPS bearings

C.M.


(4) Substructure piersγ=Σmpier/Mdeck = .05, typical

Stiff rotational springs in x & ydirections, typicalat base of piers

Q1Q2

Q3 Q4

~

88

Configuration 5: Elevated Single-span with Un-symmetric Stiffness

Configuration 6: Elevated Single-span with Eccentric Stiffness


x

y

(4) FP bearings

C.M.


(4) Substructure piers



Symmetric pier bracing in x & y direction, optional

γ=Σmpier/Mdeck = .05, typ.~ Single Span Elevated Deck Σmpier/Mdeck = 0.10 condition shown~

x

y

(4) FP bearings

C.M.





Pier bracing added in x & ydirection, as shown

~

89

Configuration 7: Elevated Double-span

Configuration 8: Elevated Double-span with Eccentric Stiffness


(6) FP bearings






Q1Q2

Q3 Q4

Pier 5Pier 6

x

y

C.M.

~

(6) FP bearings


Stiff rotational springs at Q3 & Q4


Symmetric pier bracing in x & y direction at endshown only

Q1Q2

Q3 Q4

Pier 5Pier 6

x

y

C.M.

Flexible rotational springs at Q1 & Q2,x-direction


and Pier 5 & 6,x-direction

~

90

Configuration 9: Elevated Double-span with Yielding Piers 5 and 6

Configuration 10: Substructure Static Pullback Tests






Q1Q2

Q3 Q4

Pier 5Pier 6

x

y

C.M.

(6) FP bearings


Yielding rotationalspring plates inx & y-direction atPiers 5 & 6 only

(A514 - 100ksi)

(A514 - 100ksi nom.)

(A36 - 36ksi nom.)

~

Deck frame and bearings(6) Substructure piers



Q1Q2

Q3 Q4

56

x

y Yielding rotationalspring plates inx & y-direction atPiers 5 & 6 only(A36 - 36ksi nom.)

(A514 - 100ksi nom.)

(A514 - 100ksi nom.)

Typical pullback rigging:tension cable and straps,load cell, and come-alongin series at top of pier

removed, typical

91

Figure 3-7 Instrumentation plan for Configurations 1 through 3 of non-elevated bridge deck model

Q

◆ ◆ ◆

◆◆ ◆

◆◆

◆

92

Figure 3-8 Supplementary instrumentation for Configurations 4 through 9 of elevated bridge deck model

93

a. Short plate (a.k.a., stiff spring) b. Long plate (a.k.a., flexible spring)

Figure 3-9 Typical strain gauge layout at pier base rotational leaf springs




1 LVDT displacement horizontal table h1o stroke2 LVDT displacement horizontal table h2o stroke3 LVDT displacement horizontal table h3o stroke4 LVDT displacement horizontal table h4o stroke5 DCDT displacement vertical table v1o stroke6 DCDT displacement vertical table v2o stroke7 DCDT displacement vertical table v3o stroke8 DCDT displacement vertical table v4o stroke9 A acceleration horizontal table h1-2 acc.10 A acceleration horizontal table h3-4 acc.11 A acceleration horizontal table h4-1 acc.12 A acceleration horizontal table h2-3 acc.13 A acceleration vertical table 1v acc.14 A acceleration vertical table 2v acc.15 A acceleration vertical table 3v acc.16 A acceleration vertical table 4v acc.17 LC force horiz. shear x load cell - bearing Q118 LC force horiz. shear y load cell - bearing Q119 LC force moment y load cell - bearing Q120 LC force moment x load cell - bearing Q121 LC force axial load cell - bearing Q122 LC force horiz. shear x load cell - bearing Q223 LC force horiz. shear y load cell - bearing Q224 LC force moment y load cell - bearing Q225 LC force moment x load cell - bearing Q2

C L

3”


Plan View (N.T.S.)

C L

3” 12”


SG, top ofplate only

Plan View (N.T.S.)

94

26 LC force axial load cell - bearing Q227 LC force horiz. shear x load cell - bearing Q328 LC force horiz. shear y load cell - bearing Q329 LC force moment y load cell - bearing Q330 LC force moment x load cell - bearing Q331 LC force axial load cell - bearing Q332 LC force horiz. shear x load cell - bearing Q433 LC force horiz. shear y load cell - bearing Q434 LC force moment y load cell - bearing Q435 LC force moment x load cell - bearing Q436 LC force axial load cell - bearing Q437 A acceleration horiz. y mass - acc138 A acceleration horiz. y mass- acc239 A acceleration horiz. x mass - acc340 A acceleration horiz. x mass - acc441 A acceleration horiz. x mass - acc542 A acceleration horiz. y deck frame - acc643 A acceleration horiz. y deck frame - acc744 A acceleration horiz. x deck frame - acc845 A acceleration horiz. x deck frame - acc946 A acceleration horiz. x deck frame - acc1047 A acceleration horiz. x bearing Q1 - acc1148 A acceleration horiz. y bearing Q1 - acc1249 A acceleration vertical bearing Q1 - acc1350 A acceleration horiz. x bearing Q2 - acc1451 A acceleration horiz. y bearing Q2 - acc1552 A acceleration vertical bearing Q2 - acc1653 A acceleration horiz. x bearing Q3 -acc1754 A acceleration horiz. y bearing Q3 - acc1855 A acceleration vertical bearing Q3 - acc1956 A acceleration horiz. x bearing Q4 - acc2057 A acceleration horiz. y bearing Q4 - acc2158 A acceleration vertical bearing Q4 - acc2259 LP displacement horiz. y deck frame - LP160 LP displacement horiz. y deck frame - LP261 LP displacement horiz. x deck frame - LP362 LP displacement horiz. x deck frame - LP463 LP displacement horiz. x deck frame - LP564 LP displacement horiz. y bearing Q1 - LP665 LP displacement horiz. x bearing Q1 - LP766 LP displacement horiz. y bearing Q2 - LP867 LP displacement horiz. x bearing Q2 - LP968 LP displacement horiz. y bearing Q3 - LP10


95

69 LP displacement horiz. x bearing Q3 - LP1170 LP displacement horiz. y bearing Q4 - LP1271 LP displacement horiz. x bearing Q4 - LP1372 LP displacement horiz. y table - LP1473 LP displacement horiz. x table - LP1574 LP displacement horiz. x table - LP16







1 thru 58 |Same as Configurations 1 through 3 - non-elevated Bridge Deck, see Table 3-3 above

59 A acceleration horizontal Pier Q1 x - acc2360 A acceleration horizontal Pier Q1 y - acc2461 A acceleration horizontal Pier Q2 x - acc2562 A acceleration horizontal Pier Q2 y - acc2663 A acceleration horizontal Pier Q3 x - acc27

Channel 63 converted to uni-axial force readings for Config. 10 pullback tests, as follows63 LC force horizontal Uni-Axial load cell64 A acceleration horizontal Pier Q3 y - acc2865 A acceleration horizontal Pier Q4 x - acc2966 A acceleration horizontal Pier Q4 y - acc30




70 SG linear strain curvature y Pier Q1 - XOP71 SG linear strain horiz. x Pier Q1 - XFT72 SG linear strain horiz. x Pier Q1 - XTH73 SG linear strain curvature x Pier Q1 - YOP74 SG linear strain horiz. y Pier Q1 - YFT


96

75 SG linear strain horiz. y Pier Q1 - YTH76 SG linear strain curvature y Pier Q2- XOP77 SG linear strain horiz. x Pier Q2- XFT78 SG linear strain horiz. x Pier Q2 - XTH79 SG linear strain curvature x Pier Q2 - YOP80 SG linear strain horiz. y Pier Q2 - YFT81 SG linear strain horiz. y Pier Q2 - YTH82 SG linear strain curvature y Pier Q3 - XOP83 SG linear strain horiz. x Pier Q3 - XFT84 SG linear strain horiz. x Pier Q3 - XTH85 SG linear strain curvature x Pier Q3 - YOP86 SG linear strain horiz. y Pier Q3 - YFT87 SG linear strain horiz. y Pier Q3 - YTH88 SG linear strain curvature y Pier Q4- XOP89 SG linear strain horiz. x Pier Q4- XFT90 SG linear strain horiz. x Pier Q4 - XTH91 SG linear strain curvature x Pier Q4 - YOP92 SG linear strain horiz. y Pier Q4 - YFT93 SG linear strain horiz. y Pier Q4 - YTH94 SG curvature curvature y Pier 5 - XOP95 SG linear strain horiz. x Pier 5 - XFT96 SG linear strain horiz. x Pier 5 - XTH97 SG curvature curvature x Pier 5 - YOP98 SG linear strain horiz. y Pier 5 - YFT99 SG linear strain horiz. y Pier 5 - YTH100 SG curvature curvature y Pier 6- XOP101 SG linear strain horiz. x Pier 6- XFT102 SG linear strain horiz. x Pier 6 - XTH103 SG curvature curvature x Pier 6 - YOP104 SG linear strain horiz. y Pier 6 - YFT105 SG linear strain horiz. y Pier 6 - YTH106 LC force horiz. shear x load cell - bearing 5107 LC force horiz. shear y load cell - bearing 5108 LC force moment y load cell - bearing 5109 LC force moment x load cell - bearing 5110 LC force axial load cell - bearing 5111 LC force horiz. shear x load cell - bearing 6112 LC force horiz. shear y load cell - bearing 6113 LC force moment y load cell - bearing 6


97

114 LC force moment x load cell - bearing 6115 LC force axial load cell - bearing 6

Channels 116 through 128 skipped129 LP displacement horiz. y deck frame - LP1130 LP displacement horiz. y deck frame - LP2131 LP displacement horiz. x deck frame - LP3132 LP displacement horiz. x deck frame - LP4133 LP displacement horiz. x deck frame - LP5134 LP displacement horiz. y bearing Q1 - LP6135 LP displacement horiz. x bearing Q1 - LP7136 LP displacement horiz. y bearing Q2 - LP8137 LP displacement horiz. x bearing Q2 - LP9138 LP displacement horiz. y bearing Q3 - LP10139 LP displacement horiz. x bearing Q3 - LP11140 LP displacement horiz. y bearing Q4 - LP12141 LP displacement horiz. x bearing Q4 - LP13142 LP displacement horiz. y table - LP14143 LP displacement horiz. x table - LP15144 LP displacement horiz. x table - LP16145 LP displacement horiz. x Pier Q1 - LP17146 LP displacement horiz. x Pier Q1 - LP18147 LP displacement horiz. y Pier Q2 - LP19148 LP displacement horiz. x Pier Q2 - LP20149 LP displacement horiz. y Pier Q3 - LP21150 LP displacement horiz. x Pier Q3 - LP22151 LP displacement horiz. y Pier Q4 - LP23152 LP displacement horiz. x Pier Q4 - LP24153 LP displacement horiz. y bearing 5- LP25154 LP displacement horiz. x bearing 5 - LP26155 LP displacement horiz. y bearing 6 - LP27156 LP displacement horiz. x bearing 6 - LP28157 LP displacement horiz. y Pier 5 - LP29158 LP displacement horiz. x Pier 5 - LP30159 LP displacement horiz. y Pier 6 - LP31160 LP displacement horiz. x Pier 6 - LP32

The following Channels were relocated for Pier 5 & 6 acceleration readingsnote: acc6, acc7, acc8, acc10 to Pier 5x, Pier 5y, Pier 6x, Pier6y respectively

42 A acceleration horiz. x Pier 5 - acc31


98

table motions produced are referred to as LA13_14, NF01_02, and LS17c_18c and

utilized associated similitude length scale factors of lr = 1/2, 1/5, and 1/5, respectively, to

enforce conformance with table limits. Each table motion consists of a bi-directional pair

of time-histories oriented orthogonally along the x- and y- axes of the shake table (e.g.,

LA13_14 consists of the scaled x-direction LA13 and y-direction LA14 components). The

scaled near-fault motion (i.e., NF01_02) includes its vertical component as well. Motions

could be run separately, as uni-directional (x- or y-direction) components, or as bi-

directional pairs. In the case of the NF01_02 motion, the vertical component could be

added or omitted from the table time-history. Response spectra for the LA13_14,

NF01_02, and LS17c_18c shake table motions are shown in Figure 3-10 for comparison.

In addition, to study the effect of similitude scaling on the bridge specimen (see Section

3.1.1 discussion) these motions were processed with alternate length scale factors. The

effect of varying similitude length scale (lr) on the response spectra of three of these table

motion components is illustrated in Figure 3-11.

Near-fault ground motions have been shown to place significant demand on structures.

The impulsive and sometimes long-period content of these motions may produce velocity

and displacement demands significantly exceeding the design criteria provided by code

specified spectrum compatible ground motions. These unique earthquake time-histories

may be modeled with sufficient accuracy (for structures responding at or near the peak in

43 A acceleration horix. y Pier 5 - acc3244 A acceleration horiz. x Pier 5 - acc3346 A acceleration horix. y Pier 5 - acc34


99

the earthquake’s response spectra) as simple pulse motions [Krawinkler and Alavi, 1998].

On this basis, a suite of bi-directional, pure pulse motions were developed for these

studies with pure cosine and sine waves selected to model fault-normal and fault-parallel

displacement histories, respectively. Figure 3-12 illustrates time-histories for these

simulated motion pairs at their peak displacement amplitude of 5 inches (the limit of the

shake table apparatus). Fault-normal pulse durations of Tp = 1 and 2 seconds were selected

providing a range bracketing characteristic periods (i.e., Tiso) of the test bearings (see

Table 3-1). The duration of associated fault-parallel motions were taken as 2/3, 1, and 3/2

of Tp to evaluate the coupling effect between these orthogonal components. Table 3-5

shows amplitude and duration parameters for the entire suite of idealized pulse motions

developed from these parameters. These synthetic motions are representative of suggested

pulse parameters for actual near-fault earthquake records as recommended in [Krawinkler

and Alavi, 1998] at a similitude length scale factor of lr = 1/4.

To estimate friction coefficients for the various Friction Pendulum slider composites

provided by the manufacturer, a suite of sinusoidal acceleration time-histories were also

developed. These motions were designed to impose a steady-state response of varying

displacement and velocity amplitudes on the test specimen. Results from these tests were

utilized to estimate instantaneous friction values for the different composites. Figure 3-13

illustrates a representative acceleration time-history of these sinusoidal motions. Signal

amplitude and frequencies used to develop the various test motions are in Table 3-6.

100

a. Displacement spectra: LA13_14 b. Total acceleration spectra: LA13_14

c. Displacement spectra: NF01_02 d. Total acceleration spectra: NF01_02

e. Displacement spectra: LS17c_18c f. Total acceleration spectra: LS17c_18c

Figure 3-10 Displacement and total acceleration spectra for LA13_14, NF01_02, and LS17c_18c table motions, ζ=5%

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

la13la14

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

3

T (sec)

A to

tal (

g)

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

nf01nf02

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

3

T (sec)

A to

tal (

g)

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

ls17cls18c

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

3

T (sec)

A to

tal (

g)

101

a. Displacement spectra: LA13 b. Total acceleration spectra: LA13

c. Displacement spectra: NF01 d. Total acceleration spectra: NF01

e. Displacement spectra: LS17c f. Total acceleration spectra: LS17c

Figure 3-11 Response spectra of LA13, NF01, and LS17c table motions with varying length scale, ζ = 5%

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

SF=2SF=4

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

T (sec)

A to

tal (

g)

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

SF=3SF=5

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

T (sec)

A to

tal (

g)

0 0.5 1 1.5 20

5

10

15

T (sec)

D (

in)

SF=3SF=5

0 0.5 1 1.5 20

0.5

1

1.5

2

2.5

T (sec)

A to

tal (

g)

102

3.3 Pre-Test Analysis of Bridge Deck Model

An analysis matrix, developed to include all proposed model configurations (see Section

3.1.5), was prepared for pre-test analytical evaluations of the test specimen. Pre-test

analysis assumed properties for specimen substructure components based upon their

design values (see Section 3.1.4). For systems employing FP bearings, nominal friction

coefficients of 0.04, 0.06, 0.09, and 0.12 were considered to bound the range of expected

values. FP bearing stiffness was assumed based upon the design value of the dish radius, R

a. Fault-normal b. Fault-parallel

Figure 3-12 Idealized near-fault pulse displacement histories

Table 3-5 Characteristics of near-fault pulse table motions at full amplitude

Label Fault-orientation Tp (sec) Dgmax (in.) Vgmax (in./sec) Agmax (g)

nfpu

lse1

_

normal 1 5 31.4159 0.5108

parallel 0.667 5 47.1003 1.1483

parallel 1 5 31.4159 0.5108

parallel 1.5 5 20.9440 0.2270

nfpu

lse2

_

normal 2 5 15.7080 0.1277

parallel 1.33 5 23.6210 0.2888

parallel 2 5 15.7080 0.1277

parallel 3 5 10.4720 0.0568

0 0.5 1−5

0

5

time/Tp

Dg (

g)

0 0.5 1−5

0

5

time/Tp

Dg (

in)

103

(see Section 3.1.4). LR bearing properties were averaged from available characterization

test data. Analyses were performed utilizing the generalized multi-degree-of-freedom, bi-

directional bridge model developed previously (see Section 2.3.3). The three earthquake

time-history motions (i.e., LA13_14, NF01_02, and LS17c_18c) and the suite of idealized

near-fault pulse motions (i.e., nfpulse1_ and nfpulse2_, see Table 3-5) were considered

over a range of amplitudes.

Summarized results from these pre-test analyses are included in Appendix B. These

results were examined to establish table motion amplitudes which would likely produce

specimen response near the limits of bearing and component displacement capacities.

Figure 3-13 Representative acceleration time-history of sin signal table motion

Table 3-6 Sin signal characteristics for characterization of FP slider µ values

Ao (g) f (Hz) Do (in) Vo (in/sec)

.130 1.714 0.4331 4.6643

.145 2.285 0.2718 3.9024

.111 2.285 0.20807 2.98740

.097 2.856 0.11639 2.08867

.104 6.000 0.02827 1.0659

0 2 4 6 8 10 12 14−0.15

−0.1

−0.05

0

0.05

0.1

0.15

time (sec)

acc

ele

ratio

n (

g)

104

Trial span settings for test motions were then selected to provide 50% and 100% of these

peak amplitudes. Final peak table span settings were tailored during simulation tests to

produce model response near capacity, with final test sequences run at 50% and 100% of

these final settings.

3.4 Experimental Results for Configuration 1 (Non-elevated, Symmetric Mass)




x-y orbit plots of the global displacement and global force, and x- & y-direction bearing

hysteresis (for quadrant 1, a.k.a., Q1) are presented as a summary of the test specimen

response. For uni-directional tests, x-direction displacement and force histories are

presented in lieu of orbits and y-direction bearing hysteresis at quadrant Q1 is omitted.

Selected results are shown only for the maximum table span setting run for each

individual test motion sequence.

The Configuration 1 specimen was tested utilizing the three selected earthquake table

motions and the suite of idealized near fault pulses (see Section 3.2). Tests were performed

at two or more amplitudes (or span settings) and bi-directional motion pairs were run

simultaneously and as separate x- & y-direction inputs. Sinusoidal bearing

characterization tests were also performed to estimate friction coefficients for FP bearing

slider composites.

For earthquake history tests, representative results for both uni-directional and bi-

directional tests are presented in Figures D-1 through D-19. These results are presented for

105

tests with FP, HDR, and LR bearings. Figure D-1 and D-5 illustrate response to the

LA13_14 earthquake table motion. Figure D-1 and D-2 compare response of a system

utilizing FP bearings with type 1 and 2 PTFE slider composite to the x-direction LA13

component. As seen, the FP bearings exhibit broad stable hysteretic response. The type 2

slider composite has a higher friction coefficient (see Section 3.14) resulting in a lower

peak displacement response. Figure D-3 shows the effect of the bi-directional LA13_14

table motion on the FP bearing system. As seen in the displacement orbit (see Figure D-

3(a)), this motion produces response oriented strongly along a 45 degree line to the xy-

axes. This behavior suggests strong coupling in the bearing response, as interaction with

the bearing yield surface is encountered. As expected, the x-direction bearing response at

quadrant Q1 (see Figure D-3(c)) is effected by this coupling, as the force response of the

bearing is reduced (compared to the bearing response without the LA14 component seen

in Figure D-2(c)), particularly in the x- direction induced by the strong coupled response

in the y- direction. Figure D-4 illustrates the response of a HDR bearing system to this

same LA13_14 table motion at a reduced span setting. As seen in the figure, the HDR

bearings exhibit a significant stiffening behavior, particularly in the first cycle excursions

to increasingly larger strain. The result of this behavior is increased force output.

However, reduced displacement response would be expected as a result. Figure D-5

illustrates the response of a system utilizing LR bearings to the LA13 component motion

at the same span setting as the previous test. As seen in the figure, the LR bearings exhibit

broad stable hysteretic response through numerous cycles (see Figure D-5(c)).

Comparisons between bearing characteristics (i.e., Qd, ku and kd) can be made by

106

examining the bearing hysteretic plots shown in these figures (see Table 3-1 and Figure 3-

1).

Figure D-6 through D-13 illustrate response for the FP, HDR, and LR bearing systems to

the LS17c_18c and the NF01_02 table motions. Comparisons between bearing behavior

are similar to the previous test results. The LS17c_18c motion produces relatively regular

harmonic response in the specimen, as noted in the displacement histories of Figure D-

6(a), D-7(a), and D-10(a). This soft-soil motion is rich in long period content (see Figure

3-10) near to the period characteristics of these test bearings, with harmonic response as

the expected result. Bi-directional coupling effects in the LS17c_18c motion are not

significant due to its strong orientation along its x-component. The near-fault NF01_02

motion has an impulsive fault-parallel (x-direction) and fault-normal (y-direction)

component which are inherently strongly coupled. This motion produces strongly coupled

behavior in the specimen as seen in the displacement orbits of Figure D-11 and D-12.

For scaled motion tests, representative results are presented in Figures D-14 through D-19.

These tests were performed utilizing FP bearings with type 1 PTFE composite. As seen in

the plots, the table motions at larger scale factors (i.e., lr) produce specimen response of

larger magnitude and longer period content (compare for example Figure D-14 to Figure

D-15). This difference in specimen behavior is to be expected, as the same motion at a

larger length scale factor produces typically larger spectral response (as seen in Figure 3-

11), particularly in the longer period range representative of the characteristic isolation

periods of the test bearings (see Table 3-1).

107

For near fault pulse tests, representative results are presented in Figures D-20 through D-

28. These tests were performed utilizing FP bearings with type 2 PTFE composite and LR

bearings. As seen in the plots, bearing response to these pure pulse motions are smooth

and cyclic. Bearing response is also strongly influence by coupling behavior induced by

the ground motion signal phasing of fault normal and parallel components. These results

show that fault parallel component pulses with durations of 2/3, 1, and 1.5 times the fault

normal pulse duration (Tp) produce decreasingly smaller coupled response, respectively.

In particular, this is noted in the decreasing bearing response along the direction of the

fault parallel component.

Figures D-29 & D-30 present results of sinusoidal bearing characterization tests for FP

bearings with type 1 and 2 PTFE composite sliders, respectively. Section 3.14 presents a

summary of these and all other FP slider characterization tests.

3.5 Experimental Results for Configuration 2 (Non-elevated, Eccentric Mass)




one or two summary figures are presented. The first figure presents three plots of the

global displacement time-history, global force time-history, and the x-direction bearing

hysteresis (for quadrant Q1). For systems with mass eccentricity, the second figure

presents a summary of torsional response including: a comparison of hysteretic response

and displacement time-histories for the East and West ends of the bridge specimen (in the

direction of motion input orthogonal to the direction of mass eccentricity), and a time-

108

history of deck rotation.

The Configuration 2 specimen was tested utilizing the x-direction component of the three


settings) with mass eccentricities in the transverse direction of 0, 5, and 10 percent of the

overall bridge length. Tests were performed utilizing FP bearings (with type 2 PTFE

composite sliders) and LR bearings. Selected results are presented only for tests run at

their maximum amplitude. The figures illustrate the tendency for an increase in system

rotational response and bearing displacements at one end of the span as mass eccentricity

is increased (i.e., e/L increases from 0 to 0.05 to 0.10). This is made apparent by

comparing rotation histories and bearing hysteretic plots at the west and east ends of the

specimen for the same ground motion input with e/L = 0.05 and 0.10. For the case e/L = 0,

negligible system rotation was exhibited, as expected.

3.6 Experimental Results for Configuration 3 (Non-elevated, Eccentric Stiffness)

Table C-3 in Appendix C presents a complete log of shake-table tests performed on



two summary figures are presented. The first figure presents the x-direction global

displacement time-history, global force time-history, and bearing hysteresis (for quadrant

Q1). The second figure presents a summary of torsional response including: a comparison

of hysteretic response and displacement time-histories for the East and West ends of the

bridge specimen (in the direction of motion input orthogonal to the direction of stiffness

eccentricity), and a time-history of global deck rotation.

109

The Configuration 3 specimen was tested transversely utilizing the x-direction component

of the three earthquake table motions. Isolation system stiffness eccentricity was provided

by installing LR bearings and HDR bearings on the West and East ends of the bridge,

respectively. Selected results are presented only for tests run at their maximum amplitude.

Test results for this sequence are indicative of the stiffness eccentricity inherent in the

isolation system. Characteristic bearing hysteretic properties are compared in Table 3.1. It

is apparent the LR test bearings have higher initial stiffness (i.e., larger ku) and somewhat

higher characteristic strength (Qu). On the other hand, the HDR test bearings have larger

second-slope stiffness (kd) than the LR bearings at similar peak strains. The HDR bearings

also exhibit scragging effects resulting in increased force output in initial cycles. These

variations produce eccentricity in system strength, and first- and second-slope stiffness,

resulting in torsional behavior. The tendency in the test sequences is for the HDR bearings

(at one end of the span) to exhibit smaller displacement response than the LR bearings (on

the opposite end). This is presumably the result of the larger second-slope stiffness of the

HDR bearings, even though the LR bearings are initially stronger and stiffer. For the first

test cycles to large strain (see Figure D-61 and 62 for the LA13_14 test motion), HDR

bearings exhibited first cycle scragging effects resulting in force output larger than the LR

bearings. In later tests after scragging had occurred and insufficient time had been allotted

between tests to allow recovery in the HDR bearings, LR and HDR bearings exhibited

similar force output (despite the larger displacement response of the LR bearings).

110

3.7 Experimental Results for Configuration 4 (Elevated Single-span with Sym-metric Stiffness)




two summary figures are presented. The first figure presents the global displacement and

force orbits and x- & y-direction bearing hysteresis (for quadrant Q1). The second figure


pier hysteretic response in the x- & y-direction (at quadrant Q1).

The Configuration 4 specimen was tested utilizing the three selected earthquake time-

histories. Tests were performed at two or more amplitudes (or span settings) and bi-

directional motion pairs were run simultaneously and as separate x- & y-direction inputs.

Tests were performed utilizing FP bearings with type 4 PTFE composite. Selected results

are presented only for the bi-directional tests run at their maximum table span setting.

Similar to the Configuration 1 specimen test results (see Section 3.4), these results

indicate that response is influenced strongly by ground motion characteristics, including

frequency content and directional orientation. The LA13_14 motion has significantly

higher frequency content (see Section 3.2) and is strongly oriented along a 45 degree axis

to the x-y direction. This results in specimen response with significant coupling in the x-

and y-directions and underlying cycling of higher frequency (see Figure D-67 and D-68).

The LS17c_18c motion contains lower frequency components (see Section 3.2) and is

oriented strongly along the x-direction. This results in specimen response with strong

orientation along the x-direction with little coupled response and low frequency harmonic

111

cycling (see Figure D-69 and D-70). The NF01_02 motion is an impulsive near-fault event

with strong coupled x- and y-components (see Section 3.2). The vertical component

included in this ground motion history also contains significant amplitude accelerations.

This results in similar response to that of the LA13_14 motion, with less underlying high

frequency contribution, and added fluctuations in the bearing friction force component of

the hysteretic response induced by the vertical acceleration contributions (see Figure D-71

and D-72).

In addition, system attributes play an important contribution to the response as well.

Substructure pier response is essentially linear in these tests, with only slight pinching

near the origin, as seen in the hysteretic plots (see Figure D-68, D-70, and D-72 (e) and

(f)). This pinching behavior is discussed further in Section 3.13. On the other hand,

bearing response is essentially bilinear (see Figure D-67, D-69, and D-71 (c) and (d)), with

pinching of varying magnitude. This pinching response may be seen as the result of

several factors: bi-directional coupling in the bearing yield surface, vertical load

fluctuations caused by overturning effects and/or vertical acceleration input (see NF01_02

test results), or vertical load redistributions resulting from kinematic shortening of pier

assemblies. The latter phenomenon is discussed in further detail in subsequent evaluations

in Chapter 4. Finally, it is noted that total specimen displacement response (at the deck

level) is distributed between the isolation bearings and the pier substructure assemblies in

proportion to their flexibilities. Each of these components resist nearly equal force

transmission (varying slightly as the result of substructure mass contributions) with total

displacement response being the sum of the two component displacement contributions in

series.

112

3.8 Experimental Results for Configuration 5 (Elevated Single-span with Un-symmetric Stiffness)



are presented in Appendix D as Figures D-73 through D-98.

The Configuration 5 specimen was tested with and without bracing utilizing the three


settings) and bi-directional motion pairs were run simultaneously and as separate x- and y-

direction inputs. The specimen was also tested with uni-directional components of the

earthquake table motions at alternate length scale factors (i.e., lr), similar to the

Configuration 1 tests (see Section 3.4). The NF01_02 motion was run with and without its

vertical signal. Tests were performed utilizing FP bearings with type 3 and 4 PTFE

composites. Selected results are presented only for tests run at their maximum table span

setting.

For the selected earthquake history tests shown in Figure D-73 through D-92, two

summary figures are presented. The first figure presents the global displacement and force

orbits and x- and y-direction bearing hysteresis (for quadrant Q1). The second figure


pier hysteretic response in the x- and y-direction (at quadrant Q1). The characteristics of

specimen response to the different ground motion histories is similar to the Configuration

4 tests (see Section 3.7 above), with variations in ground motion frequency content and

directionality strongly influencing specimen behavior. For braced specimen tests,

substructure pier response is again nearly linear and notably stiffer than the unbraced

113

condition (compare Figure D-74 (e) and (f) to Figure D-76 (e) and (f)). For the unbraced

specimen, this configuration has larger pier stiffness in the x-direction compared to the y-

direction. This property can be readily observed from the response figures (see, for

example, Figure D-76 (e) and (f) and Section 3.1.4 and 3.1.5). Substructure stiffness

characterizations for these configurations are presented in Section 3.13 below. Foremost,

these results indicate that the effect of increasing substructure flexibility is an increase in

both global and substructure peak displacements. It appears, however, that bearing

displacement response may be somewhat larger, smaller, or relatively unaffected by this

variation (compare for example Figure D-74 to D-76). Further, FP bearing type 3 slider

composite is shown to have a lower friction value than the type 4 PTFE composite (see

Section 3.14 below), resulting in larger global and bearing displacement response when

these sliders are installed (compare Figure D-82 to D-84 (a) and (b)). Finally, it is noted

that bi-directional substructure stiffness is uncoupled in this specimen design. The effect

of this unequal x- and y-direction substructure stiffness (most notably in the unbraced

Configuration 5 specimen) is to skew the systems’ bi-directional yield surface by

producing unequal yield displacements in these directions. This effects the character of

global system hysteretic coupling, which can be seen by comparing Configuration 4 and

Configuration 5 hysteretic response for the same input motion (compare Figures D-72 and

D-86 (a) and (b)).

For uni-directional scaled motion tests in Figure D-93 through D-98, x-direction

displacement and force histories are presented in lieu of orbits, y-direction bearing

hysteresis at quadrant Q1 is omitted, and the plots showing local global, bearing, and pier

response distribution are omitted. The results of these tests are similar to the scaled motion

114

tests of Configuration 1 (see Section 3.4), with the table motions at larger length scale

factors (i.e., lr) producing specimen response of larger magnitude and longer period

content (compare Figure D-94 and D-97). This difference in behavior is similar to before,

as the isolated specimen will experience larger spectral response for the same motion at a

larger length scale (as seen in Figure 3-11).

3.9 Experimental Results for Configuration 6 (Elevated Single-span with Eccen-tric Stiffness)




tests, four summary figures are presented. The first figure presents the global

displacement and force orbits and x- and y-direction bearing hysteresis (for quadrant Q1).


plots of overall, bearing, and pier hysteretic response in the x- and y-direction at quadrant

Q1 and Q4, respectively. The fourth figure presents a summary of torsional response

including: a comparison of hysteretic response and displacement time-histories for the

East and West ends of the bridge specimen (in the direction of motion input, orthogonal to

the direction of stiffness eccentricity), and a time-history of global deck rotation.





utilizing FP bearings in two configurations. The first slider configuration utilized type 3

115

composite on all four bearings. The second slider configuration utilized type 4 and 3

composite on the East (unbraced) and West (braced) end of the bridge, respectively, in an

effort to counteract the torsional effects of the substructure stiffness eccentricity. Selected

results are presented only for bi-directional tests run at their maximum amplitude.

The torsional response induced by the specimen’s substructure stiffness eccentricity is

apparent from these test results, with global x-direction displacement (orthogonal to

stiffness eccentricity) larger on the unbraced East end relative to the braced West end (see

Figure D-102, D-106, D-110, D-114, D-118, and D-122). On the other hand, global

displacements in the y-direction are similar on the East and West end of the specimen with

displacement compatibility enforced by the longitudinally rigid deck frame (compare, for

example, plot (b) of Figure D-100 vs. D-101 and Figure D-116 vs. D-117). Bearing

displacements in the x-direction are not, however, systematically larger or smaller on the

unbraced end compared to the braced end of the specimen. This implies that shear

response is similar on either end of the specimen span. Torsional response is then mainly

due to kinematic rotation of the substructure about a vertical axis caused by the difference

in displacement between the flexible unbraced piers on the East end relative to the braced

West end piers. Bearing displacements in the y-direction, however, are always larger on

the braced West end of the specimen due to redistribution enforced by displacement

compatibility imposed by the longitudinally rigid deck frame (compare, for example, plots

(c) and (d) of Figure D-100 vs. D-101 and Figure D-116 vs. D-117).

It is also evident from these results that the alternate slider configuration, with lower

friction FP type 3 PTFE sliders installed on the braced end of the specimen and higher

116

friction type 4 sliders installed on the unbraced end, has a tendency to reduce rotational

response due to the underlying substructure stiffness eccentricity. This is illustrated by

comparing torsional response to the same input motion for the system utilizing the

uniform slider configuration and this alternate un-symmetric slider configuration

(compare Figure D-102 vs. D-106, Figure D-110 vs. D-114, and Figure D-118 vs. D-122).

3.10 Experimental Results for Configuration 7 (Elevated Double-span)













utilizing FP bearings in four configurations. The first two utilized type 4 and 5 PTFE

composite sliders, respectively, uniformly on all four bearings. The third system utilized

type 4 composite sliders on the East and West braced ends of the span and type 5

composite sliders on the flexible central piers (i.e., Piers 5 and 6) in an effort to reduce

117

shear transfer to these components. The final system utilized type 5 composite sliders on

the East and West braced ends and type 4 composite sliders on the central piers (Piers 5

and 6) in an effort to balance shear forces across all bent lines. Selected results are

presented only for bi-directional tests run at their maximum amplitude.

The effect of ground motion characteristics on specimen response in these tests is similar

to the Configuration 1, 4 and 5 test sequences. As discussed previously, input frequency

content and directionality playing a significant role in determining system dynamic

behavior. See previous discussions in Section 3.4 and 3.7.

The Configuration 7 specimen represents a simple two-span bridge overcrossing, with

relatively rigid end abutments and a flexible central pier bent. It is notable from these

results that this specimen exhibits similar global response as the Configuration 1 and

braced Configuration 5 test specimens when subjected to the same input motion (compare,

for example, Figure D-3, D-73, and D-132 for the LA13_14 input motion). This behavior

would be expected, as each of these configurations has a similar relationship of total mass

to global force-deformation behavior, with Configuration 7 having only the added

complexity of a flexible central pier bent modifying the hysteretic behavior of this

component.

The typical pattern of displacement and force distribution behavior in this specimen can

be seen by comparing hysteretic behavior at the end abutment location (i.e., Q1) to

response at the central pier bent (i.e., Q5) (see for example Figure D-124 and D-125 for

the LA13_14 input motion). The tendency is for bearings at the braced end of the

specimen to contribute nearly all of the total displacement demand at these locations.

118

Larger shear demand is also attracted to the braced abutment ends. The central pier

locations experience similar total displacement demands as end bents, due to displacement

compatibility enforced by the in-plane flexural rigidity of the deck frame. However, the

flexible piers at these locations provide a significant contribution to the total displacement

demand with bearings contributing the remainder. However, kinematic shortening of these

central pier assemblies (which occurs through rigid body rotation of these elements as the

tip of the pier is displaced) effects loss of axial force at these locations. This phenomenon

results in loss of FP bearing stiffness whose friction and pendulum stiffness components

are axial load dependant. This results in a reduction in shear force transmission at these

locations. This behavior is evident by noting the hysteretic pinching which occurs in

bearing response at these locations (see, for example, Figure D-125 and D-128 (c) and

(d)). The implications of this kinematic effect are discussed in further detail subsequently

in Chapter 4.

Bearing characterization tests show that the type 4 and 5 slider composites exhibited

similar friction values during these tests (see Section 3.14). Consequently, it is evident that

due to this similarity the effect of different slider configurations on system response

characteristics was minimal during these test sequences (compare, for example, Figure D-

123, D-126, D-129, and D-132).

3.11 Experimental Results for Configuration 8 (Elevated Double-span with Eccen-tric Stiffness)




119

tests, five summary figures are presented. The first figure presents the global displacement

and force orbits and x- & y-direction bearing hysteresis (for quadrant Q1). The second,

third and fourth figures present summaries of local substructure response including plots

of overall, bearing, and pier hysteretic response in the x- & y-direction at quadrant Q1,

Pier 5, and quadrant Q4, respectively. The fifth figure presents a summary of torsional

response including: a comparison of hysteretic response and displacement time-histories

for the East and West ends of the bridge specimen (in the direction of motion input,

orthogonal to the direction of stiffness eccentricity), and a time-history of global deck

rotation.





utilizing FP bearings with type 5 composite. Selected results are presented only for bi-

directional tests run at their maximum amplitude.

The pattern of torsional response induced by the specimen’s substructure stiffness

eccentricity is similar to the Configuration 6 test sequences. Global x-direction

displacement (orthogonal to stiffness eccentricity) increases along the span moving from

the braced West end to the unbraced central piers to the unbraced East end (see, for

example, Figure D-160, D-161, D-162 and D-163 for the LA13_14 input motion). As

before, global displacements in the y-direction are similar on the East and West end of the

specimen with displacement compatibility enforced by the longitudinally rigid deck

120

frame. Bearing displacements in the x-direction are similar implying similar shear

response along the specimen span. Torsional response is then mainly due to kinematic

rotation of the substructure about a vertical axis caused by the difference in displacement

between the piers along the span. Bearing displacements in the y-direction, however, are

always larger on the braced end of the specimen due to pier stiffness at this location and

displacement compatibility imposed by the longitudinally rigid deck frame.

3.12 Experimental Results for Configuration 9 (Elevated Double-span with Yield-ing Piers 5 and 6)













utilizing FP bearings. Type 5 PTFE composite sliders were utilized on the East and West

braced ends of the specimen while type 4 sliders were utilized on the central piers (i.e.,

Piers 5 and 6). The vertical load was also balanced such that approximately 5/8ths was

121

supported by Piers 5 and 6 (consistent with the loading of a continuous girder span) in an

effort to attract shear forces to this location.

The Configuration 9 specimen configuration is essentially identical to Configuration 7,

with the addition of weaker A36 leaf springs installed in the central piers. Several test

sequences run at increasing span settings were performed on this specimen with limited

yielding occurring in central piers (see, for example, pier hysteresis in Figure D-179 (e)

and (f)) Consequently, pier bracing at the end abutments was removed in an effort to

balance shear behavior across the specimen increasing demands at the yielding central

piers. In this unbraced configuration, span settings were again increased to maximize

ductility demand at the central piers (see pier hysteresis in Figure D-181 and D-184 (e)

and (f)). Peak ductility demand of approximately 1.5-2 were achieved in the final tests

sequence with no apparent reduction in the isolated performance of the system (see Figure

D-182, D-183, and D-184).

3.13 Experimental Results for Configuration 10 (Substructure Static Pullback Tests)

Table C-10 in Appendix C presents a complete log of static pull-back tests performed on

Configuration 10 of the isolated bridge specimen. Selected experimental results from

these tests are presented in Appendix D as Figure D-185. Braced frame stiffness

characterization was computed from dynamic test data (see, for example, Figure D-178 (e)

and (f)). Table 3-7 below tabulates substructure pier assembly properties computed from

these test results.

122

Substructure properties in Table 3-7 determined from these characterization tests can be

compared to the previous computed design values shown in Table 3-2. As seen, a number

of discrepancies between the design and test values are apparent. First of all, stiffness

values are generally lower than computed design values. Actual braced pier stiffness is

slightly less than computed (i.e., 26.7 kips/inch computed vs. 22-24 kips/inch tested). This

reduction is apparently due to friction slip of bolted end connections. In addition, actual

stiffnesses of short and long plate pier assemblies are lower than computed (i.e., 4 kips/

inch computed vs. 2.4-3 kips/inch tested for the short plate assembly and 1.8 kips/inch

computed vs. 1.6-1.7 kips/inch tested for the long plate assembly). These stiffness

reductions are likely due to gapping in the mechanical assemblies which occurred at

movements near the displacement origin. Further, assumed slider elevations were slightly

lower than as-built elevations in design computations. This led to a higher computation of

effective lateral stiffness in design due to the displacement transformation at the lower

assumed slider height. Finally, it is seen that strength values for short and long plate pier

assemblies were higher than computed (i.e, 2.9 kips design value vs. 4 kips tested value).

This effect was certainly the result of material strength above the assumed nominal value

(i.e., 36 ksi for A36 material).

3.14 Experimental characterization data for FP slider composites

A number of tests were performed during the shake table simulation sequences to evaluate

the friction coefficients of the five FP bearing slider composites provided by the

manufacturer. A log of these bearing characterization tests are shown in Table C-1, C-4,

C-5, and C-7 of Appendix C.

123

The shear response, V(u), of a spherical sliding FP bearing responding uni-directionally

can be represented by the following,

(14)

where µ is the friction coefficient of the slider-dish interface, u the uni-directional

displacement, the velocity, N the normal force (possibly varying in time), and R the

bearing radius [Constantinou et al. 1998]. It has been shown that the friction coefficient of

a PTFE slider moving across a stainless steel surface increases with velocity up to a

threshold value [Constantinou et al. 1990] (see Figure 3-14 below). Bearing

characterization tests were performed by applying test signals to the bridge specimen

which produced essentially harmonic bearing response through a range of velocities (see

Figure 3-13 and Table 3-6). Bearing shear response data from these tests were processed

Table 3-7 Substructure pier assembly properties computed from characterization data

Configuration1 Material Ksub (kips/in) α2 yield point 3

(kips)Fy 4

(kips)Tsub

5(sec)

Braced x- 22 0.27

Braced y- 24 0.26

Short Plate A514 2.4-2.6 0.83-0.80

Long Plate A514 1.6-1.7 0.99-1.02

Short Plate A36 2.4-3 .04-.167 4 4.75-5 0.83-0.74

Long Plate A36 1.6-1.7 .23-.25 4 4.6 0.99-1.02

1. see Section 3.1.4

2. strain hardening ratio, see Figure D-185

3. first point of yield onset, see Figure D-185

4. vertex of bilinear idealization, see Figure D-185

5. based upon tributary mass; see Chapter 2, Equation 3

V u( ) µN u·( )sgn NR----u+=

u·

124

to remove the stiffness contribution (i.e., N/R) and then divided by the time varying

normal force to determine the friction coefficient of each slider composite. Mean friction

coefficient results for each of the slider composite types plotted as a function of velocity

are shown in Figure 3-15. Error bars are included in the plots indicating one standard

deviation statistical scatter.

As seen from these results, friction coefficients ranged from lowest to highest for the type

1, 3, 4, 5, and 2 slider composites, respectively. These composite types (in their virgin

condition) exhibited peak average friction coefficient values of approximately 0.05,

0.0575, 0.08, 0.095, and 0.103, respectively. The type 1 slider composite exhibited a

significant increase in friction coefficient after 57 test signals (see Figure 3-15(a)). The

type 4 slider composite exhibited a similar increase in friction response, after 231 tests

were performed (see Figure 3-15(d)). Slider composite type 5, on the other hand, exhibited

a slight reduction in average friction response after 48 tests (see Figure 3-15(e)).

Figure 3-15(f) illustrates the hysteretic friction response of bearings using the type 5 slider

composite subjected to the LS17c table motion time-history. It is apparent from this figure

that higher friction response (approaching 13 percent) is exhibited in the initial

displacement cycle of the bearing. This behavior may indicate an initial “stick”

phenomenon in the slider interface not noted in the previous characterization tests which

were not processed to near zero velocity. As seen in the remaining response, however,

friction of near 10 percent is exhibited during peak displacement cycles. This is consistent

with the bearing characterization data (see Figure 3-15(e) for values of the peak average

friction coefficient plus one standard deviation). As displacement cycles subside in the

125

remaining hysteretic response, friction approaches 8%, consistent with the lower velocity

response for this composite (see Figure 3-15(e)). This is consistent with reported behavior

due to rate effects (see Figure 3-14) [Constantinou et al. 1990].

Figure 3-14 PTFE slider composite behavior as a function of velocity

velocity

fric

tion

coef

ficie

nt

126

a. Type 1 b. Type 2

c. Type 3 d. Type 4

e. Type 5 f. Type 5 friction hysteresis: LS17c

Figure 3-15 Characterization of FP bearing slider composites

0 1 2 3 40.03

0.04

0.05

0.06

0.07

0.08

0.09

velocity (in/sec)

fric

tion

coef

ficie

nt

virgin

after 57 tests

0 1 2 30.06

0.07

0.08

0.09

0.1

0.11

0.12

velocity (in/sec)

fric

tion

coef

ficie

nt

virgin

0 1 2 3 40.045

0.05

0.055

0.06

0.065

0.07

velocity (in/sec)

fric

tion

co

effic

ien

t

virgin

0 1 2 3 40.04

0.06

0.08

0.1

0.12

velocity (in/sec)

fric

tion

coef

ficie

nt

virgin

after 231tests

0 1 2 3 4 5

0.08

0.09

0.1

0.11

velocity (in/sec)

fric

tion

coef

ficie

nt

virgin

after 48tests

−3−2.5−2−1.5−1−0.5 0 0.5 1 1.5 2 2.5 3−0.15

−0.1

−0.05

0

0.05

0.1

0.15

displacement (in)

fric

tion

co

effic

ien

t

127

4 Evaluation of Bridge Model Test Data

4.1 Introduction

In this chapter, test results reported in Chapter 3 are compared with each other and with

results of simplified analyses. These comparisions are used to assess the basic

observations made for the pilot analytical studies examined in Chapter 2. Results are

presented in this chapter regarding the influence of

1. Substructure flexibility

2. Isolator strength and second-slope flexibility

3. Substructure mass

4. Number of components of horizontal excitation

5. Substructure strength

6. Superstructure mass eccentricity

7. Substructure stiffness eccentricity

8. Type of simplified analytical model used

9. Ground motion characteristics

10. Vertical components of ground motion

11. Displacemenet restraint

12. Bearing wear

128

These comnparisions focus on the particular features incorporated in the test specimens. A

more comprehensive and refined analytical assessment of many of these variables are

considered in Chapter 5.

4.2 Influence of Substructure Flexibility

Configurations 1, 4, and 5 allow comparison of isolated bridge systems having different

substructure flexibility. These configurations were tested with very similar FP bearings.

ALthough two different PTFE slider composites were used, they had similar strength

characteristics. Three earthquake table motions (i.e., LA13_14, NF01_02, and

LS17c_18c) were utilized for these tests. x- and y-direction components were run

individually and together. The specimens tested have substructure periods as shown in

Table 4-1. Based upon the processing criteria used, similitude length scale factors of 2, 5,

and 5 are associated with the LA13_14, NF01_01, and LS17c_18c ground motions,

respectively.

Table 4-1 Nominal substructure properties for bridge model Configuration 1, 4, and 5

Σ Ksub1 Tsub2

Config. PTFE slider3 x-direction y-direction x-direction y-direction

1 type 2 INF4 INF4 0 sec4 0 sec4

55 type 4 96 kip/in 96 kip/in 0.26 sec 0.26 sec

4 type 4 10.4 kip/in 10.4 kip/in 0.8 sec 0.8 sec

5 type 4 10.4 kip/in 6.6 kip/in 0.8 sec 1.0 sec

1. Computed using component stiffness characterized from test data, see Table 3-7.

2. Tsub computed for a deck weight of 65 kips, see Chapter 2, Equation 3.

3. See Appendix C and Figure 3-15

4. Assumed substructure properties for bridge deck mounted directly to simulator.

5. With cross-bracing added at bent frames.

129

The data collected from these test sequences are compared with each other and with the

results of elastic and inelastic analyses to assess the effect of substructure flexibility on the

response of an isolated bridge system. Analyses of isolated bridge systems with

mechanical properties similar to those in the bridge deck specimen were carried out using

recorded table motions. System displacement and acceleration response data taken from

the bridge specimen tests are compared to results from the comparative analyses in

Figures 4-1 through 4-6. The analysis results are presented in terms of spectra computed

for a wide range of substructure periods. The elastic analysis shown corresponds to a

single-degree-of-freedom system with a period corresponding to Tsub and ζ = 5%. A two-

degree-of-freedom nonlinear model is considered for the nonlinear analysis.

As seen in Figures 4-1 through 4-3, peak deck displacements for isolated systems with

very rigid substructures are much larger than those of similar non-isolated elastic systems

(i.e., a system with equal substructure stiffness). Nearly all the displacement in the isolated

system occurring in the isolator component. For isolated systems with relatively large

substructure flexibility, isolator displacement is roughly similar to that of a similar non-

isolated system. For these more flexible systems, total displacement is shared between the

isolation bearings and the substructure component. The proportion of displacement

contributed by the substructure increases as the flexibility of the substructure increases.

Peak deck displacement is thus seen to increase with moderately increasing substructure

flexibility. Peak isolator displacement is constant or increases slightly with substructure

flexibility.

130

As seen in Figures 4-4 through 4-6, measured peak deck acceleration response for these

isolated systems (i.e., Cdeck) is relatively constant over a range of substructure flexibility.

Similar trends are observed from the nonlinear analyses. This might be expected due to the

similar isolator or displacements. On the other hand, computed deck acceleration

undulates greatly for similar elastic systems.

The advantage of isolation is thus not always significant force reduction, but damage

control in substructure elements and more consistent response, that is less sensitivity to

structural characteristics such as stiffness. Damage control is achieved as the isolated

system endures significant displacement demands (taken in a large portion through the

isolation bearings) without yielding or damage to substructure and superstructure

components. Conventional bridges, on the other hand are designed to endure significant

substructure yielding damage in a design basis event.

It should also be noted that the trends seen here compare closely with those reported in

preliminary analyses (see Section 2.3.5.1).

4.3 Influence of Isolator Second-Slope Flexibility and Strength

Several test configurations allow evaluation of the effect of variations in isolation system

properties on system response. Configuration 1 was tested with LR and HDR bearings.

Configuration 1, 4 and 5 were tested with two different strength FP bearings. Each

configuration was tested with both uni-directional components and bi-directional motion

pairs of the three earthquake table motions (i.e., LA13_14, NF01_02, and LS17c_18c) and

several near fault pulse motions (i.e., nfpulse).

131

(a) Model displacement: LA13 (b) Displacement spectra: LA13

(c) Model displacement: NF01 (d) Displacement spectra: NF01

(e) Model displacement: LS17c (f) Displacement spectra: LS17cFigure 4-1 Test data for FP type 2 and 4 vs. nonlinear spectra for Cyiso = 0.09, ground motions at highest span settings.

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

deck isolator

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

deck isolator D elastic, 5% damping

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

132

Mechanical properties of the test bearings were characterized via three sources: the in-

plane test machine, bi-directional pseudo-static shake table testing, and harmonic

characterization tests performed on the bridge specimen [Fenves, 1998]. Characteristic

bearing hysteretic properties taken from these sources are summarized in Table 4-2 (see

also Figure 3-15). Accompanying Figure 4-7 illustrates the hysteretic response of these

three test bearings for the LA13 and LS17c motion sequences applied to Configuration 1.

Isolator displacement and base shear response from the Configuration 1 test sequences are

summarized in Table 4-3 below. These results illustrate the effect of varying isolation

system characteristics on peak response of an isolated bridge with an essentially rigid

substructure (as for Configuration 1). As seen in Table 4-3, displacement response tends to

be sensitive to strength and systematically increases as strength decreases (compare

displacement response for bearings with nearly equal first- and second-slope stiffness

characteristics (i.e., FP type 2 vs. FP type 1 or LR vs. HDR bearings). Base shear

(g) Model displacement: averaged (h) Displacement spectra: averagedFigure 4-1 Test data for FP type 2 and 4 vs. nonlinear spectra for Cyiso = 0.09, ground motions at highest span settings.

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

133

(a) Model displacement: LA14 (b) Displacement spectra: LA14

(c) Model displacement: NF02 (d) Displacement spectra: NF02

(e) Model displacement: LS18c (f) Displacement spectra: LS18cFigure 4-2 Test data for FP type 2 and 4 vs. nonlinear spectra for Cyiso = 0.09, ground motions at highest span settings.

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

deck isolator

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)


0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

134

response, on the other hand, consistently increases for bearings with higher second-slope

stiffness (compare FP vs. LR and HDR bearings). This result is indicative of the larger

second-slope stiffness of the LR and HDR bearings compared to the FP bearings, which

produces marked increases in shear response at larger displacements. In addition, the

stiffening behavior of the HDR bearings caused by scragging effects at peak (virgin)

response cycles results in further increases in shear response above those experienced by

the LR bearings (at similar displacement amplitudes). Notably, this force effect (in the

HDR bearings) does not appear to markedly change displacement response. FP bearings,

while experiencing initially higher shear response at low amplitude (due to their larger Qd

values), have markedly lower second-slope stiffness (i.e., kd) and experience relatively

lower shear response at larger amplitudes when compared to the stiffer LR and HDR

bearings. It should be noted, that these effects are indicative of the variation in

characteristic hysteretic properties represented by these three test bearings (i.e., Qd, ku,

and kd) and not necessarily their bearing type (i.e., FP, LR, and HDR). In other words,


0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

135

(a) Model displacement: LA13_14 (b) Displacement spectra: LA13_14

(c) Model displacement: NF01_02 (d) Displacement spectra: NF01_02

(e) Model displacement: LS17c_18c (f) Displacement spectra: LS17c_18cFigure 4-3 Test data for FP type 2 and 4 vs. nonlinear spectra for Cyiso = 0.09, ground motions at highest span settings.

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

deck isolator

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)


0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

136

these bearings could be re-designed such that their hysteretic properties were interchanged

(i.e., FP bearings given higher second-slope stiffness than the LR and HDR bearings, for

example) and the variation in response would be reflective.

Isolator displacement and base shear response from the Configuration 5 test sequences are

summarized in Tables 4-4 and 4-5 below. These results illustrate the effect of varying

isolation system strength on peak response of an isolated bridge having a flexible

substructure (i.e., Tsub = 0.25, 0.8 and 1.0 seconds). As seen in Table 4-4 and 4-5,

displacement response tends to be sensitive to strength and systematically increase as

strength decreases (see the mean displacement ratio in Table 4-4 and Table 4-5). Further,

these results indicate that strength effects displacements more as the substructure becomes

more rigid (compare Table 4-4 mean displacement ratio to Table 4-5). This appears more

apparent for larger amplitude motions. Force response, on the other hand, is less sensitive


0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

137

(a) Model acceleration: LA13 (b) Deck acceleration spectra: LA13

(c) Model deck acceleration: NF01 (d) Deck acceleration spectra: NF01

(e) Model deck acceleration: LS17c (f) Deck acceleration spectra: LS17cFigure 4-4 Test data for FP type 2 and 4 vs. nonlinear spectra for Cyiso = 0.09, ground motions at highest span settings.

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

isolated deck elastic, 5% damping

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

138

to variations in characteristic isolator strength (see the mean force ratio in Table 4-4 and

Table 4-5).

These results compare closely with those of preliminary analyses reported in Chapter 2

(see Section 2.3.5.2). It was shown there that isolator characteristic strength had a

profound effect on deck and isolator displacements, particularly for systems with the most

rigid substructures (and/or first-slope stiffness). With regard to force response, it was

shown that peak force changes mostly with variations in substructure flexibility and

isolator second-slope flexibility, but insignificantly with isolator strength (for larger

amplitude responses).

4.4 Influence of Substructure Mass

Configuration 5 of the bridge deck model also allowed comparison of an isolated bridge

system incorporating substructures with a range of contributing mass. This configuration

utilized FP bearings with two different strength PTFE slider composites and was tested

(g) Model deck acceleration: averaged (h) Deck acceleration spectra: averagedFigure 4-4 Test data for FP type 2 and 4 vs. nonlinear spectra for Cyiso = 0.09, ground motions at highest span settings.

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

139

(a) Model deck acceleration: LA14 (b) Deck acceleration spectra: LA14

(c) Model deck acceleration: NF02 (d) Deck acceleration spectra: NF02

(e) Model deck acceleration: LS18c (f) Deck acceleration spectra: LS18cFigure 4-5 Test data for FP type 2 and 4 vs. nonlinear spectra for Cyiso = 0.09, ground motions at highest span settings.

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)


0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

140

with both uni-directional and bi-directional ground motions on the earthquake simulator.

The three earthquake table motions (i.e., LA13_14, NF01_02, and LS17c_18c) were

utilized. The Configuration 5 specimen targeted three nominal substructure mass

proportions, defined as γ = msub/Mdeck (where msub and Mdeck are the total substructure

and deck mass, respectively), equal to approximately 0, 5, and 10 percent. Arrangements

of substructure mass, FP slider composites (see Chapter 3, Section 3.14), and bracing

arrangements used for these test sequences are outlined in Table 4-6 below.

Bridge specimen peak response results from these tests are plotted in Figure 4-8 through

4-11. As seen in the first three of these figures, results indicate the effect of substructure

mass on system response is most prominent on substructure displacement, less so on

isolator displacement, and least so on total deck displacement on average (see Figure 4-8,

4-9, and 4-10 (g) and (h)). In any event, the difference in average response rarely exceeded

10 percent in these tests for a variation of substructure mass of γ = 0 to 10 percent. The

(g) Model deck acceleration: averaged (h) Deck acceleration spectra: averagedFigure 4-5 Test data for FP type 2 and 4 vs. nonlinear spectra for Cyiso = 0.09, ground motions at highest span settings.

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

141

(a) Model deck acceleration: LA13_14 (b) Deck acceleration spectra: LA13_14

(c) Model deck acceleration: NF01_02 (d) Deck acceleration spectra: NF01_02

(e) Model deck acceleration: LS17c_18c (f) Deck acceleration spectra: LS17c_18cFigure 4-6 Test data for FP type 2 and 4 vs. nonlinear spectra for Cyiso = 0.09, ground motions at highest span settings.

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)


0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

142

tendency is for the average response to increase as the substructure mass ratio, γ,

increases. This implies base shear response is highest for a system with the most

substructure mass (resulting directly from the higher substructure displacement demand),

as expected. On the other hand, Figure 4-11 indicates that the substructure mass ratio has a

(g) Model deck acceleration: averaged (h) Deck acceleration spectra: averaged

Table 4-2 Summary idealized bilinear hysteretic characteristics for test bearings

Bearing Type Qd ku kd Tiso1

FP, type 1 0.81 kips INF2 0.54 kips/in 1.75sec

FP, type 2 1.63 kips INF2 0.54 kips/in 1.75 sec



LR3 1.46 kips 5 kip/in 0.9 kips/in 1.36 sec

HDR4 0.81 kips 3.5 kip/in 1 kips/in 1.29 sec

1. Tiso computed for a deck weight of 65 kips applied uniformly to four bearings.

2. Assumed theoretical initial stiffness for friction device.

3. LR bearing properties at γ = 200%, see Table 3-1.

4. HDR bearing properties at γ = 250%, see Table 3-1.

Figure 4-6 Test data for FP type 2 and 4 vs. nonlinear spectra for Cyiso = 0.09, ground motions at highest span settings.

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

0 0.2 0.4 0.6 0.8 10

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

Tsub

(sec)

Cdeck

(g

)

143

Table 4-3 Influence of isolation system on Configuration 1 bridge model response

Response Displacement (in.) Base shear (g)

Bearing Type FP,type 1

FP,type 2 LRB HDR FP,

type 1FP,

type 2 LRB HDRH

alf a

mpl

itude

LA13 0.969 2.082 2.052 0.122 0.198 0.237NF01 0.885 1.724 1.733 0.126 0.181 0.190NF02 0.552 0.919 0.114 0.120

NF01_02c 1.361 1.792 0.138 0.198LS17c 0.176 1.448 2.019 0.126 0.161 0.193LS18c 0.063 0.791 0.106 0.105

LS17c_18c 0.247 2.177 0.125 0.197

Full

ampl

itude

NF01 3.500 3.521 0.229 0.423NF02 2.635 2.633 0.193 0.268

NF01_02c 4.560 3.817 0.276 0.497LS17c 2.051 1.729 4.176 0.149 0.161 0.498LS18c 0.662 1.482 0.129 0.163

LS17c_18c 1.672 4.288 0.149 0.480nfpulse1x 6.536 6.001 0.287 0.283

nfpulse0667 2.161 1.469 0.140 0.148nfpulse1_0667 6.862 6.370 0.289 0.290

nfpulse1x 6.496 5.836 0.287 0.285nfpulse1y 0.621 0.121 0.087 0.103

nfpulse1_1 6.389 5.725 0.268 0.269

nfpulse1x 6.372 5.809 0.285 0.286nfpulse15 0.033 0.063 0.059 0.080

nfpulse1_15 6.229 5.749 0.282 0.282nfpulse2_1333 3.696 2.386 4.148 0.216 0.181 0.310

nfpulse2_2 2.497 1.198 3.159 0.164 0.139 0.264

nfpulse2_3 2.605 1.01 3.294 0.172 0.143 0.274

144

relatively insignificant effect on isolator shear force response. Due to the low second-

slope stiffness of these devices, this is to be expected since increased isolator displacement

results in only minor increases in force output.

Table 4-4 Influence of isolation system on response of Configuration 5 bridge deck model, braced, Tsub = 0.25 seconds (nominal)

Response Displacement (in.) Isolator total shear force (kips)

Bearing Type FP, type 3 FP, type 4 ratio FP, type 3 FP, type 4 ratio

Hal

f am

plitu

de

LA13 1.118 0.986 1.134 6.515 8.223 0.792LA14 1.206 1.535 0.786 7.179 9.091 0.790

LA13_14 1.513 2.039 0.742 7.182 9.092 0.790NF01 1.398 1.247 1.121 7.904 8.461 0.934NF02 0.880 0.640 1.375 7.432 7.562 0.983

NF01_02c 1.826 1.659 1.101 7.435 7.566 0.983LS17c 0.484 0.424 1.142 7.702 8.131 0.947LS18c 0.210 0.186 1.134 7.190 7.492 0.960

LS17c_18c 0.415 0.429 0.967 7.191 7.500 0.959

mean = 1.056 1.002

C.O.V. = 0.186 0.083

Full

ampl

itude

LA13 3.363 2.795 1.203 12.434 11.915 1.044LA14 4.442 3.929 1.130 14.820 14.779 1.003

LA13_14 5.001 4.481 1.116 14.836 14.813 1.002NF01 4.314 4.056 1.064 15.784 16.179 0.976NF02 3.355 2.811 1.193 12.666 12.309 1.029

NF01_02c 5.363 5.035 1.065 12.672 12.321 1.028

LS17c 2.325 2.178 1.068 10.510 10.776 0.975

LS18c 0.961 0.898 1.070 7.897 8.574 0.921LS17c_18c 2.401 2.162 1.110 7.899 8.575 0.921

mean = 1.113 0.989

C.O.V. = 0.049 0.045

145

These results compare closely with those of preliminary analyses reported in Section

2.3.5.3. These preliminary analyses indicated that systems with 10 percent or less of the

total mass lumped in the substructure, exhibit approximately 90 percent or more of the

Table 4-5 Influence of isolation system on response of Configuration 5 bridge deck model, Tsub = 0.8 and 1.0 seconds (nominal) in x- and y-direction, respectively.

Response Displacement (in.) Isolator total shear force (kips)

Bearing Type FP, type 3 FP, type 4 ratio FP, type 3 FP, type 4 ratio

Hal

f am

plitu

de

LA13 2.067 2.332 0.886 8.660 9.924 0.873LA14 2.693 2.503 1.076 10.168 10.813 0.940

LA13_14 2.865 2.684 1.067 10.171 10.816 0.940NF01 2.178 2.071 1.051 9.066 10.452 0.867NF02 1.801 1.584 1.137 8.070 8.971 0.900

NF01_02c 2.669 2.595 1.029 8.072 8.971 0.900LS17c 2.0319 2.050 0.991 8.0004 10.191 0.785LS18c 0.582 0.594 0.980 3.222 3.438 0.937

LS17c_18c 2.042 2.137 0.956 3.222 3.438 0.937

mean = 1.019 0.898

C.O.V. = 0.073 0.057

Full

ampl

itude

LA13 4.322 4.048 1.068 12.820 13.119 0.977LA14 6.623 6.454 1.026 16.735 16.870 0.992

LA13_14 6.947 6.701 1.037 16.742 16.875 0.992NF01 5.472 5.218 1.049 15.387 15.652 0.983NF02 5.994 4.679 1.281 15.460 13.535 1.142

NF01_02c 7.945 6.606 1.203 15.472 13.552 1.142

LS17c 3.294 3.513 0.937 11.688 12.781 0.914

LS18c 1.167 1.175 0.993 6.502 6.921 0.939LS17c_18c 3.395 3.410 0.995 6.505 6.922 0.940

mean = 1.065 1.002

C.O.V. = 0.102 0.083

146

(a) FPS, type 2 bearings: LA13 (b) FPS, type 2 bearings: LS17c

(c) LRB bearings: LA13 (d) LRB bearings: LS17c

(e) HDR bearings: LA13 (f) HDR bearings: LS17c

Figure 4-7 Test data for bearing hysteresis, ground motions at highest span settings.

−2 −1 0 1 2−15

−10

−5

0

5

10

15

X − Displ. (in.)

X −

She

ar (

kips

)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)


−2 −1 0 1 2

−10

−5

0

5

10

X − Displ. (in.)

X −

She

ar (

kips

)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

−2 −1 0 1 2−15

−10

−5

0

5

10

15

X − Displ. (in.)

X −

She

ar (

kips

)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

−2 −1 0 1 2

−10

−5

0

5

10

X − Displ. (in.)

X −

She

ar (

kips

)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

−2 −1 0 1 2−15

−10

−5

0

5

10

15

X − Displ. (in.)

X −

She

ar (

kips

)

0 0.2 0.4 0.6 0.8 10

1

2

3

4

5

6

7

8

9

Tsub

(sec)

Dis

p. (

in)

−2 −1 0 1 2

−10

−5

0

5

10

X − Displ. (in.)

X −

She

ar (

kips

)

147

global deck level peak displacement as a lumped mass (single-degree-of-freedom)

representation of the same system. This test substantiates these previous findings. More

comprehensive studies examining the effect of substructure mass on system response are

developed subsequently in Chapter 5.

4.5 Influence of Bi-directional Motions

Tests on the various configurations of the bridge deck model allowed for the evaluation of

the effect of bi-directional motions on the response of isolated bridge systems. All

configurations were tested using both the uni-directional components of each ground

motion input as well as their bi-directional pairs.

The effect of bi-directional input was evaluated by computing the following two ratios:

(15)

where Uxy is the peak vectored displacement response to the bi-directional pair and Ux and

Uy are the peak displacement responses to the individual uni-directional input of the

ground motion applied separately on the x- and y- direction, respectively; and

Table 4-6 Configuration 5: substructure mass, PTFE slider, and bracing arrangements

γ PTFE slider Cross-bracing x-direction, Tsub y-direction, Tsub

0 type 4 No .75 sec 1 sec

.05 type 4 Yes 0.25 sec .25 sec

.05 type 4 No .75 sec 1 sec

.05 type 3 Yes 0.25 sec .25 sec



Uxy mean Ux Uy,( )⁄

148

(a) LA13_14, half amplitude (b) LA13_14, full amplitude

(c) NF01_02, half amplitude (d) NF01_02, full amplitude

(e) LS17c_18c, half amplitude (f) LS17c_18c, full amplitude

Figure 4-8 Total displacement (U), test data for FP type 4, ground motions at half and full amplitude

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

2.5

3

msub

/Mdeck

U (

in) x−direction

y−direction xy−direction

0 0.02 0.04 0.06 0.08 0.10

2

4

6

8

msub

/Mdeck

U (

in)

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

2.5

3

msub

/Mdeck

U (

in)

0 0.02 0.04 0.06 0.08 0.10

2

4

6

8

msub

/Mdeck

U (

in)

x−direction y−direction xy−direction

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

2.5

3

msub

/Mdeck

U (

in)

0 0.02 0.04 0.06 0.08 0.10

2

4

6

8

msub

/Mdeck

U (

in)

149

(16)

where srss(Ux,Uy) is the SRSS (i.e., square root of the sum of the squares) combination of

Ux and Uy.

Results of these ratios are tabulated in Tables 4-7 through 4-11 below for the various

specimen configurations for systems with HDR and FP bearings. Nominal periods T0, T1,

T2, T3, and T4 represent “initial” or non-isolated specimen periods (i.e., Tsub, see Equation

3) for the Configuration 1 through 7 specimen arrangements. Nominal values were

computed utilizing substructure component stiffness characterization data (see Table 3-7).

These results indicate that the bi-directional coefficient ratios appear to generally decrease

as substructure flexibility increases, indicating that the effect of bi-directional input is

most pronounced for systems with the most rigid initial stiffnesses (see Figure 4-12). The

Uxy/mean(Ux,Uy) and Uxy/srss(Ux,Uy) coefficients cover the range of 1.4-1.8 and 0.9-1.2

(g) Mean values, half amplitude (h) Mean values, full amplitude

Figure 4-8 Total displacement (U), test data for FP type 4, ground motions at half and full amplitude

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

2.5

3

msub

/Mdeck

U (

in)

0 0.02 0.04 0.06 0.08 0.10

2

4

6

8

msub

/Mdeck

U (

in)

Uxy srss Ux Uy,( )⁄

150



(e) LS17c_18c, half amplitude (f) LS17c_18c, full amplitudeFigure 4-9 Isolator displacement (Uiso), test data for FP type 4, ground motions at half and full amplitude.

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

msub

/Mdeck

Uis

o (in

)


0 0.02 0.04 0.06 0.08 0.10

1

2

3

4

5

msub

/Mdeck

Uis

o (in

)


0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

msub

/Mdeck

Uis

o (in

)

0 0.02 0.04 0.06 0.08 0.10

1

2

3

4

5

msub

/Mdeck

Uis

o (in

)

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

msub

/Mdeck

Uis

o (in

)

0 0.02 0.04 0.06 0.08 0.10

1

2

3

4

5

msub

/Mdeck

Uis

o (in

)

151

on average, respectively, for systems with the most rigid substructures. In comparison, if

the peak response Ux and Uy were equal, a value of Uxy/mean(Ux,Uy) = 1.414 and Uxy/

srss(Ux,Uy) = 1.0 would indicate that the peak Ux and Uy displacements were occurring

simultaneously in the bi-directional response, leading to a peak Uxy equal to their vector

sum. Further in-depth studies, carried out to better establish these effects over a broader

range of ground motion inputs and bridge configurations, are presented subsequently in

Chapter 5.

4.6 Influence of Substructure Strength

Configuration 9 of the bridge deck specimen provided for an evaluation of the effect of

yielding substructure components on the response of an isolated bridge system. This

configuration utilized FP bearings with type 5 PTFE composite sliders installed on braced

end piers (abutment ends) and type 4 PTFE composite sliders installed on center, yielding

piers (see Section 3.14 for PTFE slider characterization results). Force-displacement

(g) Mean values, half amplitude (h) Mean values, full amplitudeFigure 4-9 Isolator displacement (Uiso), test data for FP type 4, ground motions at half and full amplitude.

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

msub

/Mdeck

Uis

o (in

)

0 0.02 0.04 0.06 0.08 0.10

1

2

3

4

5

msub

/Mdeck

Uis

o (in

)

152



(e) LS17c_18c, half amplitude (f) LS17c_18c, full amplitudeFigure 4-10 Substructure displacement (Usub), test data for FP type 4, ground motions at half and full amplitude.

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

msub

/Mdeck

Usu

b (in

)

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

2.5

3

msub

/Mdeck

Usu

b (in

)

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

msub

/Mdeck

Usu

b (in

)

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

2.5

3

msub

/Mdeck

Usu

b (in

)

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

msub

/Mdeck

Usu

b (in

)

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

2.5

3

msub

/Mdeck

Usu

b (in

)

153

characterizations of the high-strength end piers and yielding central piers may be

referenced from Configuration 10 test results (see Section 3.13). Axial load on the

specimen was balanced such that approximately 5/8th’s of the vertical deck weight was

supported on central piers (consistent with the load distribution of a continuous two-span

girder) to attract shear force and induce yielding in central piers. Tests with both uni-

directional and bi-directional inputs of the three earthquake table motions (i.e., LA13_14,

NF01_02, and LS17c_18c) were performed.

Representative results from these tests for global and component hysteretic response are

plotted in Figure 4-13 and 4-14 below. These plots show a pattern of force redistribution

as central piers yield and shear is transferred to outer (non-yielding) end piers. As seen in

these plots, bearing hysteresis at end piers (i.e., quadrant Q1) stiffens near peak negative

displacement indicating increase in load. Meanwhile, pier hysteresis at quadrant Q1

remains essentially linear elastic and pier hysteresis at Pier 5 (a.k.a., Q5) shows plastic

(g) Mean values, half amplitude (h) Mean values, full amplitudeFigure 4-10 Substructure displacement (Usub), test data for FP type 4, ground motions at half and full amplitude.

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

msub

/Mdeck

Usu

b (in

)

0 0.02 0.04 0.06 0.08 0.10

0.5

1

1.5

2

2.5

3

msub

/Mdeck

Usu

b (in

)

154



(e) LS17c_18c, half amplitude (f) LS17c_18c, full amplitudeFigure 4-11 Isolator shear force (V), test data for FP type 4, ground motions at half and full amplitude

0 0.02 0.04 0.06 0.08 0.10

2

4

6

8

10

12

msub

/Mdeck

V (

kips

) x−direction y−direction xy−direction

0 0.02 0.04 0.06 0.08 0.10

5

10

15

20

msub

/Mdeck

V (

kips

)


0 0.02 0.04 0.06 0.08 0.10

2

4

6

8

10

12

msub

/Mdeck

V (

kips

)

0 0.02 0.04 0.06 0.08 0.10

5

10

15

20

msub

/Mdeck

V (

kips

)

0 0.02 0.04 0.06 0.08 0.10

2

4

6

8

10

12

msub

/Mdeck

V (

kips

)

0 0.02 0.04 0.06 0.08 0.10

5

10

15

20

msub

/Mdeck

V (

kips

)

155

yielding in the negative cycle (indicated by widening of the hysteretic loop). Despite this

yielding behavior, total deck displacement history remains stable as seen in the

accompanying time-histories (see Figure 4-13 and 4-14 (b)). This indicates that isolated

(g) Mean values, half amplitude (h) Mean values, full amplitude

Table 4-7 Uxy/mean(Ux,Uy) coefficients, FP type 2 and 4, ground motions at half and full amplitude

PTFE composite type 2 type 4 type 4 type 4 type 4

Tsub, x-direc-tion 0 sec 0.25 sec 0.75 sec 0.75 sec 0.25 sec

Tsub, y-direc-tion 0 sec 0.25 sec 0.75 sec 1.0 sec 0.25 sec

Hal

f am

plitu

de LA13_14 1.490 1.618 1.056 1.110 1.110NF01_02 1.894 1.758 1.512 1.420 1.609LS17_18c 2.058 1.407 1.473 1.617 1.329Average 1.814 1.594 1.347 1.382 1.349

Full

ampl

itude LA13_14 1.378 1.333 1.236 1.276 1.215

NF01_02 1.486 1.467 1.453 1.335 1.438

LS17_18c 1.399 1.406 1.358 1.455 1.645Average 1.421 1.402 1.349 1.355 1.433

Figure 4-11 Isolator shear force (V), test data for FP type 4, ground motions at half and full amplitude

0 0.02 0.04 0.06 0.08 0.10

2

4

6

8

10

12

msub

/Mdeck

V (

kips

)

0 0.02 0.04 0.06 0.08 0.10

5

10

15

20

msub

/Mdeck

V (

kips

)

156

Table 4-8 Uxy/mean(Ux,Uy) coefficients, FP type 3 and 5, ground motions at half and full amplitude


Tsub, x-direc-tion

T0 = 0 sec T1 = 0.25 sec T2 = 0.8 sec T2 = 0.8 sec T4 = 0.30 sec

Tsub, y-direc-tion


Hal

f am

plitu

de LA13_14 1.302 1.204 1.078NF01_02 1.603 1.342LS17_18c 1.194 0.650 1.405Average 1.366 1.065 1.242

Full

ampl

itude LA13_14 1.282 1.269 1.238

NF01_02 1.399 1.386 1.404LS17_18c 1.461 1.522 1.590Average 1.380 1.392 1.411

Table 4-9 Uxy/srss(Ux,Uy) coefficients, FP type 2 and 4, ground motions at half and full amplitude


Tsub, x-direc-tion


Tsub, y-direc-tion


Hal

f am

plitu

de LA13_14 1.044 1.118 0.745 0.785 0.785NF01_02 1.305 1.183 1.048 0.995 1.115

LS17_18c 1.317 0.927 0.935 1.001 0.874

Average 1.222 1.076 0.910 0.927 0.925

Full

ampl

itude LA13_14 0.971 0.929 0.869 0.880 0.852

NF01_02 1.041 1.020 1.009 0.943 1.002

LS17_18c 0.903 0.918 0.891 0.921 1.031Average 0.972 0.956 0.923 0.914 0.962

157

performance is maintained, despite the yielding substructure contribution. In this isolated

system, isolation bearings absorb the majority of total displacement demand up to yielding

in the substructure. Beyond this point, central piers yield plastically until peak

displacement is attained. As such, only minor ductility demands are imposed on the

substructure. For a non-isolated system, nearly the entire displacement demand would be

Table 4-10 Uxy/srss(Ux,Uy) coefficients, FP type 3 and 5, ground motions at half-amplitude


Tsub, x-direc-tion


Tsub, y-direc-tion


Hal

f am

plitu

de LA13_14 0.920 0.844 0.762NF01_02 1.105 0.944LS17_18c 0.786 0.356 0.951Average 0.937 0.715 0.857

Full

ampl

itude LA13_14 0.898 0.878 0.868

NF01_02 0.982 0.979 0.983LS17_18c 0.954 0.972 0.981Average 0.944 0.943 0.944

Table 4-11 Uxy/mean(Ux,Uy) and Uxy/srss(Ux,Uy) coefficients, HDR, x-direction Tsub = y-direction Tsub = T0 = 0 sec

Uxy/mean(Ux,Uy) Uxy/srss(Ux,Uy)

Amplitude half full half full

Reco

rd

LA13_14 1.392 .979NF01_02 1.351 1.241 .913 .868

LS17_18c 1.550 1.516 1.004 .968

Average 1.431 1.378 .966 .918

158

(a) Uxy/mean(Ux,Uy): LA13_14 (b) Uxy/srss(Ux,Uy): LA13_14

(c) Uxy/mean(Ux,Uy): NF01_02 (d) Uxy/srss(Ux,Uy): NF01_02

(e) Uxy/mean(Ux,Uy): LS17c_18c (f) Uxy/srss(Ux,Uy): LS17c_18cFigure 4-12 Bi-directional coefficients computed from test data, FP type 2 and 4, ground motions at full amplitude.

0 0.2 0.4 0.6 0.81

1.1

1.2

1.3

1.4

1.5

Tsub

(sec)

Uxy

/mea

n(U

x,Uy)

0 0.2 0.4 0.6 0.80.85

0.9

0.95

1

1.05

Tsub

(sec)

Uxy

/srs

s(U

x,Uy)

0 0.2 0.4 0.6 0.81

1.1

1.2

1.3

1.4

1.5

Tsub

(sec)

Uxy

/mea

n(U

x,Uy)

0 0.2 0.4 0.6 0.80.85

0.9

0.95

1

1.05

Tsub

(sec)

Uxy

/srs

s(U

x,Uy)

0 0.2 0.4 0.6 0.81

1.1

1.2

1.3

1.4

1.5

Tsub

(sec)

Uxy

/mea

n(U

x,Uy)

0 0.2 0.4 0.6 0.80.85

0.9

0.95

1

1.05

Tsub

(sec)

Uxy

/srs

s(U

x,Uy)

159

imposed on the substructure elements, inducing significantly larger ductility demands on

these components.

Figure 4-15 illustrates the character of the force redistribution seen in these tests. This

redistribution may be visualized clearly for a static application of lateral force to a

representative isolated, elasto-perfectly-plastic, bridge pier assembly. As shown in the

figure, as load is increased to the level of yielding in the substructure pier, no further load

can be developed in the system at increasing displacement (due to the elastic-perfectly-

plastic pier response). As such, isolator displacement ceases to increase. With moderate

hardening in the yielding substructure component (i.e., α>0), a moderate increase in force

and isolation system displacement would be realized during plastic yielding of the

substructure bent (see Figure 4-15). In a system with additional redundant components (of

higher strength), additional force increases will be developed through redistribution to

these elements. Pseudo-static tests performed by Kawashima, et al [Kawashima and Shoji,

1999] validate this type of redistributive behavior. Under dynamic loading, this behavior

(g) Uxy/mean(Ux,Uy): Mean values (h) Uxy/srss(Ux,Uy): Mean valuesFigure 4-12 Bi-directional coefficients computed from test data, FP type 2 and 4, ground motions at full amplitude.

0 0.2 0.4 0.6 0.81

1.1

1.2

1.3

1.4

1.5

Tsub

(sec)

Uxy

/mea

n(U

x,Uy)

0 0.2 0.4 0.6 0.80.85

0.9

0.95

1

1.05

Tsub

(sec)

Uxy

/srs

s(U

x,Uy)

160

is further complicated by damping, higher mode contributions, axial load redistributions,

and bi-directional interactions. However, as seen in these test results, the nature of this

redistributive mechanism maintains a similar character in the dynamic response as well.

(a) Global system hysteresis (b) Global displacement history

(c) Q1 global hysteresis (d) Q5 global hysteresisFigure 4-13 Configuration 9, Test No. 23, LA13_14, Span Setting [870/0/0]

−6 −4 −2 0 2 4 6

−20

−10

0

10

20

Disp. (in)

She

ar (

kips

)

10 15 20 25

−6

−4

−2

0

2

4

6

time (sec)D

isp.

(in

)

−5 0 5

−4

−2

0

2

4

Disp. (in)

She

ar (

kips

)

−5 0 5

−4

−2

0

2

4

Disp. (in)

She

ar (

kips

)

161

4.7 Influence of Superstructure Mass Eccentricity

Configuration 2 of the bridge deck model allowed the evaluation of the effect of

superstructure mass eccentricity on the response of an isolated bridge overcrossing

(supported on an essentially rigid supporting substructure). This configuration utilized

both FP and LR bearings. Only x- direction uni-directional components of the table

motions (i.e., LA13, NF01, and LS17c) were imposed orthogonal to the direction of mass

(e) Q1 bearing hysteresis (f) Q5 bearing hysteresis

(g) Q1 pier hysteresis (h) Q5 pier hysteresisFigure 4-13 Configuration 9, Test No. 23, LA13_14, Span Setting [870/0/0]

−5 0 5

−4

−2

0

2

4

Disp. (in.)

She

ar (

kips

)

−5 0 5

−4

−2

0

2

4

Disp. (in)

She

ar (

kips

)

−5 0 5

−4

−2

0

2

4

Disp. (in)

She

ar (

kips

)

−5 0 5

−4

−2

0

2

4

Disp. (in.)

She

ar (

kips

)

162

(a) Global system hysteresis (b) Total displacement time-history

(c) Q1 global hysteresis (d) Q5 global hysteresis

(e) Q1 bearing hysteresis (f) Q5 bearing hysteresis

Figure 4-14 Configuration 9, Test No. 24, LA13_14, Span Setting [1000/0/0]

−5 0 5

−20

−10

0

10

20

Disp. (in)

She

ar (

kips

)

5 10 15 20 25−5

0

5

time (sec)

Dis

p. (

in)

−5 0 5−5

0

5

Disp. (in)

She

ar (

kips

)

−5 0 5−5

0

5

Disp. (in)

She

ar (

kips

)

−5 0 5−5

0

5

Disp. (in.)

She

ar (

kips

)

−5 0 5−5

0

5

Disp. (in)

She

ar (

kips

)

163

eccentricity. In order to avoid excessive damage to or permanent displacement of the LR

bearings, displacement amplitudes were limited to half that imposed on the FP bearings.

Table 4-12 below lists results from these tests. Results tabulated are the normalized

isolator displacement ratio, De/Do, a comparative measure between the peak isolator

displacement of a system with and without given mass eccentricity, e (i.e., De and Do,

respectively). These limited test results appear to indicate that torsional contributions to

peak isolator displacement (located at the ends of the deck span) are more prominent for

impulsive and soft-soil type motions than for far-field events (compare NF01 and LS17c

to LA13 results in Table 4-12). Further, the effect of mass eccentricity on peak isolator

response appears to be nonlinear (i.e., isolator displacements at the ends of the span do

not, in general, increase linearly with mass eccentricity).

Previous evaluations by others on isolated systems suggest FP bearings provide a re-

centering mechanism [Bozzo, et al 1989]. The friction force and lateral stiffness of an FP

(g) Q1 pier hysteresis (h) Q5 pier hysteresis

Figure 4-14 Configuration 9, Test No. 24, LA13_14, Span Setting [1000/0/0]

−5 0 5−5

0

5

Disp. (in)

She

ar (

kips

)

−5 0 5−5

0

5

Disp. (in.)

She

ar (

kips

)

164

(a) Isolated bridge deck (b) Idealized dynamic model

(c) Substructure force-displacement (d) Isolator force-displacement

(e) Total system force-displacement response at deck

Figure 4-15 Force redistribution in isolated bridge substructure assembly

di

dsub

d

Substructure bent

Isolation bearings

F

Mdeck

d

kukd

mpier

Csub

dsub

di

KsubαKsub

Qd/(1-kd/ku)Fysub

bent isolator

F

Displ.

Ksub

αKsubα=0

dysubdsub

Fysub

F’

F

Displ.

kd

dyiso di’

Fysub

Qd/(1-kd/ku)

di

F’

ku

α>0

F

K2

Fysub

Qd/(1-kd/ku)

dysub+di

F’

K1

α=oα>0

dsub+di dsub+di’ Displ.

165

bearing is proportional to the applied normal force (see Chapter 3, Equation 14). This

indicates that the center of resistance of an isolated system utilizing FP bearings will

inherently align with the systems center of mass, counteracting torsional response.

Analysis and testing in these previous studies confirmed this behavior [Bozzo et al, 1989].

The tests performed in this report using impulsive motions, however, seem to refute this

generalization. This may be due to the sensitivity of “rigid” structures to strength (see

Section 2.3.5.2 and 4.3 above). The lighter end of the specimen (which has a lower overall

friction force) may tend to displace more than the heavier (i.e., stronger) side producing

the torsional behavior seen here. Comparable results for full intensity motions are not

available for the LR bearings (whose strength is not sensitive to axial applied force) to test

this assumption. In this respect, further analysis and testing is still needed to establish

reliable trends in the response of these eccentric systems.

4.8 Influence of Substructure Stiffness Eccentricity

Configuration 6 and 8 of the bridge deck model allowed the evaluation of the effect of

substructure stiffness eccentricity on the response of an isolated bridge overcrossing.

Configuration 6 (an eccentric version of the Configuration 5 specimen) utilized FP

bearings with two arrangements of PTFE sliders: (1) a symmetric arrangement using type

Table 4-12 Effect of superstructure mass eccentricity on peak isolator displacement: isolator displacement ratio, De/Do

FP bearings (100% motions) LR bearings (50% motions)

Ground Motion e/L =0.05 e/L =0.10 e/L =0.05 e/L =0.10

LA13 1.03 1.18 0.94 1.15

NF01 1.17 1.24 1.21 1.16

LS17c 1.48 1.47 1.34 1.26

166

3 sliders uniformly; and (2) an unsymmetric arrangement arranged to counteract torsional

response (with type 3 sliders on stiffer braced end piers and type 4 sliders on flexible end

piers). Configuration 8 (an eccentric version of the Configuration 7 specimen) utilized FP

type 5 sliders uniformly. Each of these configurations was tested with both uni-directional

components and bi-directional motion pairs of the three earthquake table motions (i.e.,

LA13_14, NF01_02, and LS17c_18c).

Table 4-13 and 4-14 lists results from these tests. Results tabulated are for motion input at

peak amplitudes only. These limited results indicate that peak rotation demand in these

eccentric configurations was not a function of peak displacement amplitude, nor was it

effected significantly by bi-directional input. First of all, stiffness eccentricity in the

supporting substructure is significant in these configurations with the braced and unbraced

ends of the specimen differing in total lateral stiffness by nearly ten times. Prior to

yielding in the isolation bearings, this stiffness eccentricity induces significant torsional

response. After isolation bearings yield, stiffness eccentricity is minimal as the bearings

operate in their isolated mode on either end of the specimen. The rotation response in

these eccentric configurations is therefore due to the tendency of the system to rotate prior

to yielding in the bearings. After bearings have yielded, stiffness eccentricity is nearly

eliminated and rotation demands become insensitive to peak response. Thus, isolated

behavior suppresses torsional behavior to an extent, unlike in a standard bridge where

substructure stiffness eccentricity would impose torsional demands increasing with

response amplitude. Secondly, the orthogonal input motion in these tests imposes loading

on the longitudinal axis of the specimen where the stiffness arrangement is symmetric.

This motion input does not engage lateral-torsional coupling in the system and therefore

167

does not tend to increase rotation demands. Coupling in the yield response of the isolation

bearings imposed by the bi-directional input does not appear to effect peak rotation

significantly either, as torsional response in these systems appears to be more effected by

behavior prior to bearing yield, as just discussed.

Comparing response of the single-span specimen with and without substructure stiffness

eccentricity (i.e., Configuration 5 vs. Configuration 6, see Table 4-13), it was seen for the

case of symmetric bearing strength (i.e., type 3 FP sliders installed uniformly) that

substructure stiffness eccentricity (in the Configuration 6 specimen) tended to increase

isolator response at the ends of the specimen above both the fully braced and unbraced

Configuration 5 arrangements in all instances. This would indicate that torsional rotation

was closely phased with peak response in these test sequences, amplifying local bearing

displacement demands. An unsymmetric configuration utilizing lower and higher strength

PTFE slider composites installed in bearings located above braced and unbraced supports,

respectively, was utilized in the Configuration 6 arrangements successfully to reduce

torsional response. This result was contrary to preliminary analysis in Chapter 2 which

indicated that this type of unsymmetric arrangement of bearing strength would not

significantly reduce torsional response due to substructure stiffness eccentricity (see

Section 2.3.5.4).

Comparing response of the double-span specimen with and without substructure stiffness

eccentricity (i.e., Configuration 7 vs. Configuration 8, see Table 4-14), results indicated

that torsional response tended to increase peak total displacement response above the

Configuration 7 arrangement in all cases. Peak isolator displacement response was not as

168

systematically effected, however (sometimes increasing and sometimes decreasing),

indicating that torsional rotation was not systematically in phase with peak response in

these test sequences.

See also the results of preliminary evaluations in Chapter 2 which showed other notable

response trends due to variations in substructure stiffness eccentricity (see Section

2.3.5.4).

Table 4-13 Comparison of peak response for Configuration 5 and 6

Configuration 5 (e/L = 0) Configuration 6 (e/L = .09)

braced unbraced FP, type 3 FP, type 3 and 4

Record d di d di d di θ (rads) d di θ (rads)

LA13 3.36 3.30 4.32 3.80 4.37 4.42 .0062 4.34 4.32 .0054

NF01 4.31 4.21 5.47 4.73 5.71 5.42 .0062 5.17 4.91 .0044

LS17c 2.33 2.26 3.29 2.75 3.93 3.76 .0089 3.76 3.65 .0064

LA13_14 4.52 3.55 6.66 4.79 5.14 4.80 0.0081 5.05 4.66 0.0036

NF01_02 4.48 4.87 7.27 6.92 5.49 7.42 0.0078 5.17 6.78 0.0051

LS17c_18c 2.40 2.50 3.38 4.11 3.21 4.68 0.0078 3.27 4.56 0.0046

Table 4-14 Comparison of peak response for Configuration 7 and 8

Configuration 7 (e/L = 0) Configuration 8 (e/L = 0.10)

FP, type 5 FP, type 5

Record d di d di θ (rads)

LA13 3.49 3.37 4.01 3.38 0.0100

NF01 4.50 4.52 5.32 4.56 0.0109

LS17c 2.72 2.74 3.13 2.78 0.0119

LA13_14 4.66 3.30 5.37 3.05 0.0103

NF01_02 4.63 4.69 6.10 4.94 0.0110

LS17c_18c 2.76 2.85 3.05 2.75 0.0121

169

4.9 Response of Two-span Isolated Bridge

4.9.1 Introduction

Configuration 7 of the specimen provided for testing of a continuous double-span isolated

bridge overcrossing. This configuration utilized FP bearings with several different

arrangements of PTFE sliders. However, bearing characterization tests show that the type

4 and 5 slider composites used exhibited similar friction values during these tests (see

Section 3.14). Consequently, the effect of different slider configurations on system

response characteristics was minimal. In this configuration, the bridge deck girder

spanned continuously between braced end pier supports (abutments) and central pier

components (see Section 3.1.5). The double-span configuration was tested with both uni-

directional and bi-directional components of the LA13_14, NF01_02, and LS17c_18c

earthquake table motions.

Several noteworthy features in system response were identified from test data. These

features included a characteristic distribution of overall displacement and force demand

between local substructure and isolation bearing components. Also, local kinematic

effects in the response of the specimen’s supporting central bridge piers resulted in

variations in axial load distributions which effected overall response characteristics.

4.9.2 Characteristic distribution of force and displacement demands

Figure 4-16 and 4-17 show typical test sequence results of global and local force and

displacement response of the Configuration 7 specimen. The results in Figures 4-16 and 4-

17 (a) and (b) indicate how displacements at the end and center supports are distributed

between isolator and substructure elements. Part (c) through (h) of these figures illustrate

170

these the local hysteretic response of each isolator at these locations. As seen in the plots,

shear demands tend to concentrate toward the stiffer end (abutment) piers during the time-

history (compare Figure 4-16 (c) to (d), for example). We note that center pier-isolator

assemblies see similar total displacement demands as end pier-isolator assemblies as

displacement compatibility is enforced by the high in-plane stiffness of the bridge deck as

it displaces laterally. Therefore, since the secant stiffness of the central pier-isolator

assemblies at peak displacement is lower than at the end (abutment) pier-isolator

assemblies (for all arrangements of slider compounds used), higher shear demands

concentrate at abutment locations. Further, the displacements in the isolator and pier

components are distributed in relation to the relative flexibilities of these elements at their

respective locations. Therefore, since each location sees nearly equal total displacement

demand, a smaller proportion of the total displacement at center piers is taken in the

isolation bearings with a larger proportion taken in the more flexible center piers

(compared to isolator and pier displacement distribution at stiffer end abutment piers).

4.9.3 Local geometric effects

In the Configuration 7 specimen, end (abutment) piers were braced while center piers

could rotate on the clevis pins at their base (see Chapter 3, Section 3.1.4). During

earthquake testing flexible center piers underwent significant lateral tip displacement

(essentially a rigid body rotation). This rotational behavior caused second-order reduction

of the vertical height of these central piers. On the other hand, the shorter and stiffer

braced end (abutment) piers did not experience this effect nearly as much. because of the

high out of plane stiffness of the specemin deck, shortening in center piers resulted in a

redistribution of axial load between the center and end piers that varied according to

171

(a) Q1, abutment displacement histories (b) Q5, pier displacement histories


(e) Q1, bearing hysteresis (f) Q5, bearing hysteresisFigure 4-16 System displacement and axial load redistribution, Configuration 7, Test No.5, LA13_14, Span Setting [696/0/0]

8 9 10 11 12−5

0

5

time (sec)

Dis

p. (

in)

8 9 10 11 12−5

0

5

time (sec)

Dis

p. (

in)

global bearingpier

−2 0 2

−3

−2

−1

0

1

2

3

Disp. (in)

She

ar (

kips

)

−2 0 2

−3

−2

−1

0

1

2

3

Disp. (in)

She

ar (

kips

)

−2 0 2

−3

−2

−1

0

1

2

3

Disp. (in.)

She

ar (

kips

)

−2 0 2

−3

−2

−1

0

1

2

3

Disp. (in)

She

ar (

kips

)

172

lateral displacement. Representative axial force time-histories are shown in Figure 4-18

illustrating this effect for the LA13 motion (i.e., Test No.5, see also Figure 4-16). Figure 4-

18 (c) shows this variation in axial load by plotting the time-history of total axial load at

end and center piers, respectively, normalized by initial total load (P0). As seen in this

plot, as center piers unload (due to the axial shortening), axial load increases at abutment

piers, illustrating the load redistribution.

Since the slip-force an lateral stiffness of FP bearings are proportional to the supported

axial load (see Chapter 3, Equation 14), the axial force redistribution in these test

sequences causes a significant alteration of the characteristic force-deformation behavior

of the individual FP bearings atop the central piers. This behavior is seen clearly in the

hysteretic plots in Figure 4-16 (e) and (f) for this same test sequence. The hysteresis of

bearings at end piers (see Figure 4-16 (e)) are seen to “fatten” with increased lateral

displacement (due to increase in axial load), while the hysteresis of bearings at central

(g) Q1, substructure hysteresis (h) Q5, substructure hysteresisFigure 4-16 System displacement and axial load redistribution, Configuration 7, Test No.5, LA13_14, Span Setting [696/0/0]

−2 0 2

−3

−2

−1

0

1

2

3

Disp. (in)

She

ar (

kips

)

−2 0 2

−3

−2

−1

0

1

2

3

Disp. (in.)

She

ar (

kips

)

173

(a) Q1, abutment displacement histories (b) Q5, pier displacement histories


(e) Q1, bearing hysteresis (f) Q5, bearing hysteresisFigure 4-17 System displacement and axial load redistribution, Configuration 7, Test No.11, NF01_02, Span Setting [804/0/0]

8 9 10 11 12−5

0

5

time (sec)

Dis

p. (

in)

8 9 10 11 12−5

0

5

time (sec)

Dis

p. (

in)

global bearingpier

−4 −2 0 2 4

−6

−4

−2

0

2

4

6

Disp. (in)

She

ar (

kips

)

−4 −2 0 2 4

−6

−4

−2

0

2

4

6

Disp. (in)

She

ar (

kips

)

−4 −2 0 2 4

−6

−4

−2

0

2

4

6

Disp. (in.)

She

ar (

kips

)

−4 −2 0 2 4

−6

−4

−2

0

2

4

6

Disp. (in)

She

ar (

kips

)

174

piers (see Figure 4-16 (f)) “narrow” and soften with increasing lateral displacement (due

to loss of axial load).

Globally, total axial force (due to gravity) remains constant. Therefore, total bearing slip-

force remains essentially constant and total system strength is unchanged. Locally, the

redistribution of axial force to the end piers tends to increase peak shear demands at these

locations, imposing higher design forces on associated connections and substructure

components.

If bearings at the center and end piers had different friction coefficients, the redistribution

of axial force would increase or reduce overall system strength somewhat. For systems

with small substructure periods, this strength variation could have a significant effect on

peak displacements. Further, redistribution of axial load to the stiffer end piers produces a

system with higher overall stiffness in its isolated mode (i.e., see K3 in Figure 2-3). The

(g) Q1, substructure hysteresis (h) Q5, substructure hysteresisFigure 4-17 System displacement and axial load redistribution, Configuration 7, Test No.11, NF01_02, Span Setting [804/0/0]

−4 −2 0 2 4

−6

−4

−2

0

2

4

6

Disp. (in)

She

ar (

kips

)

−4 −2 0 2 4

−6

−4

−2

0

2

4

6

Disp. (in.)

She

ar (

kips

)

175

effect of this higher might be to increase shear force demand and decrease peak

displacements.

It is important to note that this particular axial load redistribution phenomenon appears to

be more a characteristic of the unique test setup used in this case than that of an actual

bridge. Even for bridges with high out-of-plane stiffness, reinforced concrete piers would

be expected to elongate as they displace laterally due to the opening of flexural cracks. As

a consequence, these two second-order effects may have a tendency to cancel one another

out in an actual bridge. This is an area requiring further investigation.

However, these tests point out that kinematic effects may redistribute axial load acting on

the isolators. These effects may alter the distribution of local forces. These effects are not

typically incorporated in standard design analysis methods nor included in current design

procedures [AASHTO, 1999]. The effect of axial load fluctuations on sliding systems due

to vertical ground motion input is discussed further subsequently in this chapter.

4.10 Substitute system vs. MDOF response

Post test analyses of isolated bridge systems were carried out to compare the response

predicted by two simple, single-degree-of-freedom idealizations with measured response

of the multi-degree-of-freedom bridge specimen. The braced version of Configuration 5

( seconds) with substructure mass of was utilized for this comparison.

This specimen exhibited particularly consistent hysteretic behavior from test and was

therefore considered suitable for comparison with simple, single-degree-of-freedom

models.

Tsub 0.25≈ γ 0.05≈

176

Linear substructure stiffness properties were based upon characterization test data (see

Section 3.13). Friction properties of the FP isolation bearings were selected to closely

match characterization test data as well (see Section 3.14). First-slope bearing stiffness

(ku) was assumed essentially rigid. Second-slope stiffness (kd) was computed based upon

bearing radius of curvature (R) and an imposed normal force taken as the supported deck

weight (where kd = N/R, see Chapter 3, Equation 14). From these mechanical properties,

(a) Axial load history at abutments (b) Axial load history at center piers

(c) Axial load redistribution between center piers and abutments

Figure 4-18 System axial load redistribution, Configuration 7, Test No.5, LA13_14, Span Setting [696/0/0]

3 6 9 12−80

−70

−60

time (sec)

Axi

al (

kip

s)

3 6 9 12−40

−30

−20

time (sec)

Axi

al (

kip

s)

3 6 9 120.6

0.8

1

1.2

time (sec)

Axi

al/P

0 (ki

ps)

abutment pierscenter piers

177

an idealized bilinear system model was developed for transverse and longitudinal isolated

bridge behavior (see Chapter 2, Figure 2-2). For bi-directional response, isolator behavior

was modeled as a bilinear plasticity element utilizing a circular yield function while

substructure flexibility was assumed linear and uncoupled. On the other hand, the

linearized model was only utilized for uni-directional time-history analyses. Linearized

properties were taken as the secant stiffness of the bilinear model at peak displacement

and equivalent viscous damping determined by equating the hysteretic work done by the

bilinear system (assuming symmetric hysteretic response at peak displacement) to that of

a purely viscous damping element (see Chapter 5, Section 5.4.3). Analyses of the

“substitute” bilinear and linear spring-damper systems were carried out utilizing the

ground motion inputs taken from table records of Configuration 5 test sequences.

Nonlinear analysis was performed utilizing standard procedures [Newmark, 1959].

Linearized analysis was performed utilizing standard frequency domain integration and

required an iterative procedure to converge on system spring-damper properties consistent

with computed peak displacement.

Tables 4-15 through 4-17 present force and displacement response results from the

Configuration 5 tests and the substitute system analyses for three earthquake shake table

time-histories. Results presented are the peak total, isolator, and substructure displacement

(d, di, and dsub, respectively) and lateral force response coefficient (Cs). Total

displacement and lateral force response are computed directly from analysis for the SDF

models. Local bearing and substructure response was determined (for the SDF models) by

assuming a static application of peak force applied to the bilinear isolation bearing and

substructure components, respectively (utilizing component mechanical properties from

178

the original SDF model development). This step ignores substructure damping and higher

mode contributions to localized response, dynamic information which is not included in

either of the simplified SDF models.

Tables 4-18 and 4-19 compare substitute system analytical results to test specimen data by

normalizing analysis results by specimen data. Values greater than one in this table imply

the simplified models overestimate specimen response, while values less than one imply

they underestimate them. As seen in these results, the substitute linear SDF model tended

to provide a slightly better estimate in this study of peak displacement responses when

compared to the substitute bilinear system. The linear spring-damper model overestimated

peak total, isolator and substructure displacement response in this study by approximately

3, 1, and 33 percent on average, respectively, compared to the bilinear SDF model which

overestimated these responses by 8, 6, and 50 percent on average, respectively. With

regard to force, however, the bilinear model was more accurate in this study,

overestimating lateral force response by only nearly 4 percent on average, while the linear

model underestimated force by 12 percent on average. Regardless of mean accuracy,

however, the linear SDF model provided larger coefficients of variation on average

predictions compared to the bilinear SDF model in this study. This indicates that errors on

response using the substitute linear model were larger in this study for particular ground

motion records (when compared to the substitute bilinear system), despite being more

accurate on average. The nonlinear analysis procedure does not involve iteration and can

be easily applied to bidirectional input motion.

179

It should be noted that the results of this comparison may be a function of the chosen

mechanical properties utilized to develop the SDF characterizations. Better accuracy of

either SDF model may be possible with better selection of these idealized mechanical

parameters (or a bounding procedure applied to account for their variation). However,

these results do indicate that the bilinear model tends to provide a better prediction of peak

total displacement, isolator displacement, and lateral force response for any particular

ground motion input (having exhibited reasonable average accuracy and lower

coefficients of variation on these response quantities in this study). However, the largest

errors in both of the substitute models occurred on peak substructure displacement

predictions. This indicates that neither substitute SDF model adequately accounts for the

dynamic effect of higher mode response on local substructure demands for an MDOF

isolated bridge system, particularly when these contributions are significant (note: the

Configuration 5 specimen included only a nominal substructure mass of approximately 5

percent of the deck total mass). Bilinear system models may be improved by adding a

substructure degree of freedom to account for these separate mass and damping

contributions (only slightly complicating analytical procedures). Linearized SDF models

may not readily be amenable to this form of improvement. Further in depth evaluations of

these higher mode effects on response predictions and evaluation of the accuracy of

standard linearized analytical procedures are considered in Chapter 5.

Table 4-15 Results from Configuration 5 tests

Table Motion Cyiso d di dsub CsLA13 .0575 3.36 3.30 .07 0.19LA14 .0575 4.44 4.37 0.10 0.23

LA13_14 .0575 5.00 4.93 0.11 0.24

180

4.11 Influence of Ground Motion Characteristics

The three earthquake table motions selected for these comparative studies (i.e., LA13_14,

NF01_02, and LS17c_18c) allowed evaluation of the effect of variations in ground motion

type on the response of simple isolated bridge systems. These ground motions represent

far-field, near-fault and soft-soil motions (i.e., LA13_14, NF01_02, and LS17c_18c,

respectively). Four substructure stiffness values were incorporated in the various bridge

specimen configurations (see Table 4-1). This allowed the evaluation of system response

over a broad range of spectral frequencies. For the given deck mass, corresponding

fundamental (non-isolated) system periods for these four specimen configurations are

shown in Table 4-1 as well. These values are intended to represent full scale bridge

properties at scale factors of lr = 1/2, 1/5, and 1/5 for the LA13_14, NF01_01, and

LS17c_18c table motions, respectively. These configurations were tested with FP bearings

NF01 .0575 4.31 4.21 0.13 0.24NF02 .0575 3.35 3.28 0.10 0.19

NF01_02 .0575 5.36 5.26 0.12 0.27LS17c .0575 2.33 2.26 0.08 0.16LS18c .0575 0.96 0.93 0.06 0.12

LS17c_18c .0575 2.40 2.36 0.07 0.16LA13 .08 2.80 2.75 0.08 0.18LA14 .08 3.93 3.85 0.10 0.23

LA13_14 .08 4.48 4.41 0.11 0.23NF01 .08 4.06 3.95 0.56 0.25NF02 .08 2.81 2.75 0.08 0.19

NF01_02 .08 5.04 4.91 0.13 0.27LS17c .08 2.18 2.12 0.07 0.17LS18c .08 0.90 0.87 0.06 0.13

LS17c_18c .08 2.16 2.10 0.09 0.16

Table 4-15 Results from Configuration 5 tests

181

with two different slider compounds. Each of the uni-directional components and bi-

directional motion pairs of the three earthquake table motions were utilized in the tests.

Figure 4-19 presents 5% damped elastic response spectra for each of the ground motion

table records (i.e., LA13_14, NF01_01, and LS17c_18c at scale factors of lr = 1/2, 1/5,

and 1/5, respectively). Response spectra for both uni-directional components (x- and y-

direction) and bi-directional pairs (xy-direction) are plotted. These spectra present

displacement and total acceleration response for each record superimposed for

comparison. The spectra in Figure 4-19 illustrate the variety of spectral characteristics

contained in this suite of earthquake table motions.

Table 4-16 Results from nonlinear time-history analysis utilizing a bilinear idealization

Table Motion Cyiso d di dsub CsLA13 .0575 3.63 3.51 0.14 0.17LA14 .0575 4.83 4.72 0.17 0.21

LA13_14 .0575 5.58 5.44 0.16 0.23NF01 .0575 4.46 4.31 0.17 0.20NF02 .0575 3.58 3.46 0.13 0.17

NF01_02 .0575 5.55 5.40 0.17 0.23LS17c .0575 2.96 2.86 0.13 0.15LS18c .0575 0.96 0.91 0.07 0.09

LS17c_18c .0575 3.15 3.07 0.11 0.15LA13 .08 2.72 2.60 0.14 0.17LA14 .08 4.06 3.93 0.17 0.21

LA13_14 .08 4.81 4.68 0.16 0.23NF01 .08 3.78 3.64 0.17 0.20NF02 .08 3.12 3.02 0.14 0.18

NF01_02 .08 4.75 4.62 0.17 0.22LS17c .08 2.33 2.23 0.13 0.15LS18c .08 1.07 1.00 0.09 0.11

LS17c_18c .08 2.29 2.22 0.11 0.14

182

Bridge specimen test results to these ground motion inputs indicated that response was

influenced strongly by ground motion characteristics, including frequency content and

directional orientation (see Chapter 3 for more details). Chapter 3, Section 3.4 and 3.7,

discuss details of these response characteristics. For example, the LA13_14 motion has

significant higher frequency content (see Section 3.2) and is strongly oriented along a 45

degree azimuth relative to the x-y axes. This resulted in specimen response with

significant coupling in the x- and y-directions and underlying cycling of higher frequency.

The LS17c_18c motion, on the other hand, is a soft-soil motion with long period content

near the characteristic period of the test bearings (see Section 3.2) and produced relatively

Table 4-17 Results from linearized analysis


LA13_14 .0575NF01 .0575 4.04 3.90 0.14 0.19NF02 .0575 4.79 4.64 0.14 0.21

NF01_02 .0575LS17c .0575 3.24 3.13 0.12 0.16LS18c .0575 0.73 0.67 0.05 0.08

LS17c_18c .0575LA13 .08 2.54 2.42 0.12 0.16LA14 .08 4.47 4.32 0.15 0.22

LA13_14 .08NF01 .08 3.38 3.24 0.14 0.19NF02 .08 2.47 2.37 0.11 0.16

NF01_02 .08LS17c .08 2.58 2.46 0.12 0.16LS18c .08 0.65 0.58 0.07 0.10

LS17c_18c .08

183

harmonic response in the specimen. The near-fault NF01_02 motion has impulsive,

strongly coupled, fault-normal and fault-parallel components and, therefore, produced

strongly coupled behavior in the specimen.

Further, as seen in Figure 4-19 the peaks in spectral response for the LA13_14, NF01_02,

and LS17c_18c records show unique characteristics. The response of the bridge specimen

was determined by how its fundamental, first-slope period properties correlated with the

“short” and “long” period ranges of these ground motion spectra. To illustrate this point,

Table 4-18 Comparison of bilinear system analyses to test results


LA13_14 .0575 1.115 1.104 1.523 1.008NF01 .0575 1.034 1.024 1.302 1.042NF02 .0575 1.066 1.054 1.360 1.088

NF01_02 .0575 1.034 1.026 1.443 0.965LS17c .0575 1.275 1.265 1.570 1.187LS18c .0575 1.003 0.986 1.259 1.099

LS17c_18c .0575 1.313 1.299 1.761 1.056LA13 .08 0.972 0.945 1.772 1.161LA14 .08 1.033 1.022 1.754 1.092

LA13_14 .08 1.074 1.061 1.520 1.021NF01 .08 0.932 0.921 0.304 0.988NF02 .08 1.112 1.097 1.839 1.008

NF01_02 .08 0.943 0.941 1.348 0.925LS17c .08 1.070 1.050 1.896 1.126LS18c .08 1.195 1.145 1.594 0.810

LS17c_18c .08 1.059 1.061 1.240 1.040mean 1.078 1.064 1.510 1.043C.O.V. 0.094 0.093 0.249 0.085

184

scaled versions of the x-direction component of each of three earthquake table motions

were processed for shake table simulation studies.

Figure 4-20 presents 5% damped elastic response spectra for each of the scaled ground

motion table records. These spectra present displacement and total acceleration response

for each record at two selected scale factors, superimposed for comparison (i.e., LA13_14

at lr = 1/2 and 1/4, NF01_01 at lr = 1/3 and 1/5, and LS17c_18c at lr = 1/3 and 1/5). As

seen in the figures, acceleration spectra at smaller scale factors (i.e., smaller lr, see Chapter

Table 4-19 Comparison of linearized system analyses to test results


LA13_14 .0575NF01 .0575 0.937 0.927 1.083 0.772NF02 .0575 1.427 1.414 1.510 1.088

NF01_02 .0575LS17c .0575 1.395 1.384 1.482 0.999LS18c .0575 0.756 0.727 0.939 0.657

LS17c_18c .0575LA13 .08 0.909 0.881 1.522 0.873LA14 .08 1.137 1.121 1.576 0.983

LA13_14 .08NF01 .08 0.834 0.821 0.246 0.755NF02 .08 0.880 0.860 1.375 0.838

NF01_02 .08LS17c .08 1.185 1.161 1.746 0.976LS18c .08 0.722 0.666 1.185 0.751

LS17c_18c .08mean 1.031 1.010 1.328 0.880C.O.V. 0.226 0.238 0.316 0.147

185

3, Section 3.1.1) are shifted toward the longer period range and displacement spectra are

shifted upward. Based on similitude scaling, a structure with a given fixed period will be

in a shorter period spectral range, but have a larger magnitude displacement response, for

a motion processed at a larger length scale factor. On the other hand, the same structure

will be in a longer period spectral range, but have a smaller magnitude displacement

response, to the same motion processed at the smaller length scale factor. Further, for a

given prototype (full-scale) bridge with period Tp, a model of scale lr with period Tm =

trTp (where tr = lr1/2) subjected to an earthquake record processed at the same time scale

(i.e., tr = lr1/2), will exhibit displacement response equal to dm = lrdp, where dp is the

displacement response of the prototype bridge subjected to the full scale earthquake. See

Chapter 3, Section 3.1.1 for further discussion on similitude.

Subsequent analysis of isolated bridge overcrossings were carried out to illustrate this

effect of ground motion scaling on the response of systems incorporating substructures

with a range of stiffnesses bounding those of the bridge deck specimen. These analyses

were performed utilizing the generalized multi-degree-of-freedom bridge model (see

Figure 2-4(c) and Section 2.3.3). These systems incorporated isolation bearings with

similar mechanical characteristics as the FP test bearings. Earthquake ground motions

used for the time-history simulations were taken directly from records of the scaled table

motions used in the test program (i.e., LA13 at lr = 1/2 and 1/4, NF01 at lr = 1/3 and 1/5,

and LS17c at lr = 1/3 and 1/5). Figures 4-21 presents results from these analyses with

response data taken from the bridge specimen tests to the same input motion

superimposed for comparison. As seen in the figure, these analytical results correspond

186

closely to bridge specimen test data. Figure 4-22 presents the same analytical results

carried over a broader range of system periods.

Figure 4-22 illustrates that for isolated systems with very rigid substructures (or shorter

initial elastic periods) peak total displacement is much larger than that of a similar non-

isolated (elastic) system. For systems with greater substructure (or initial) flexibility, total

peak displacement is roughly equal to that of a similar elastic system. This behavior is

characteristic of systems in these “short period” and “long period” (or equal displacement)

ranges of spectral response, as noted earlier in preliminary analytical results (see Chapter

2, Section 2.3.5.1).

The period range defining these “short” and “long” period ranges is a characteristic of the

particular ground motion input. This can be seen clearly as the motion is scaled. For

example, for the LA13 motion the “long period” (equal displacement) range appears to

begin around Tsub = 0.55 seconds for a scale factor of lr = 1/2 (see Figure 4-21 and 4-22

(a)). However, for the same motion at a length scale factor of lr = 1/4, the “long period”

range begins around Tsub = 0.35 (see Figure 4-21 (b)). In comparison, the “long period”

range begins around Tsub = greater than 1 seconds and nearly 0.9 for the NF01 and LS17c

motions at length scale factors of lr = 1/3 and 1/5, respectively. It is apparent most clearly

for the LS17c motion that the “long period” range begins near the peak in the elastic

displacement spectrum (i.e., near 1 and 0.9 seconds for lr = 1/3 and 1/5, respectively, see

Figure 4-21 and 4-22 (e)). For the NF01 motion the “long period” range also begins near

where the elastic displacement spectrum begins “leveling off” (i.e., near 1 and 0.9 seconds

for lr = 1/3 and 1/5, respectively, see Figure 4-22 (d)). On the other hand, for the LA13_14

187

motion the “long period” range begins at a point defined more roughly by where the

acceleration spectrum begins descending (i.e., near 0.55 and 0.35 seconds for lr = 1/2 and

1/4, respectively, see Figure 4-21 and 4-22 (b)).

To illustrate the variation in response characteristics expected for systems in these

different spectral ranges, consider for example an isolated bridge system with a fixed

fundamental (or initial elastic) period of Tsub = 0.4 seconds (and having the same isolation

system properties as those represented in the analytical results presented in Figure 4-21

and 4-22). This isolated system may be in the “short period” range (i.e., exceeding elastic

displacements) for the LA13 motion at lr = 1/4. For the same motion at lr = 1/2, however,

this system would be in the “long period” range (i.e., having nearly equal displacement as

an elastic system). On the other hand, this system would be in the “short period” range for

both the NF01 and LS17c motions at lr = 1/3 and 1/5. Further, an isolated bridge system

with a fixed initial elastic period of Tsub = 1 second (having the same isolation system

characteristics) would be in the “long period” (equal displacements) range for all three

motions at all length scale factors considered. Response characteristics to different ground

motion input can be evaluated similarly. For example, a system with a fixed initial elastic

period of Tsub = 0.8 seconds (an the same isolation system properties) subjected to all three

motions at their smallest length scale factors (i.e., LA13, NF01, and LS17c at lr = 1/2, 1/3,

and 1/3, respectively) would be in the “long period” range for the LA13 motion and the

“short period” range for both the NF01 and LS17c motions (compare Figure 4-21 and 4-

22 (a), (c), and (e)). This illustrates the effect of ground motion spectral characteristics

which dictate system behavior in these general frequency ranges.

188

(a) Displacement spectra: x-direction (b) Total acceleration spectra: x-direction

(c) Displacement spectra: y-direction (d) Total acceleration spectra: y-direction

(e) Displacement spectra: xy-direction (f) Total acceleration spectra: xy-directionFigure 4-19 Comparison of 5% damped elastic response spectra for LA13_14, NF01_02, LS17c_18c ground motions, peak span settings.

0 1 2 3 40

5

10

15

20

Period (sec)

D (

in)

0 1 2 3 40

0.1

0.2

0.3

0.4

Period (sec)

At (

g)

LA13 NF01 LS17c

0 1 2 3 40

5

10

15

20

Period (sec)

D (

in)

0 1 2 3 40

0.1

0.2

0.3

0.4

Period (sec)

At (

g)

LA14 NF02 LS18c

0 1 2 3 40

5

10

15

20

Period (sec)

D (

in)

0 1 2 3 40

0.1

0.2

0.3

0.4

Period (sec)

At (

g)

LA13/14 NF01/02 LS17c/18c

189

(a) Displacement spectra: LA13 (b) Total acceleration spectra: LA13

(c) Displacement spectra: NF01 (d) Total acceleration spectra: NF01

(e) Displacement spectra: LS17c (f) Total acceleration spectra: LS17cFigure 4-20 Comparison of 5% damped elastic response spectra for scaled LA13, NF01, LS17c table motions at two length scale factors, peak span settings.

0 1 2 3 40

2

4

6

8

10

12

14

Period (sec)

D (

in)

0 1 2 3 40

0.1

0.2

0.3

0.4

Period (sec)

At(

g)

lr = 2

lr = 4

0 1 2 3 40

2

4

6

8

10

12

14

Period (sec)

D (

in)

0 1 2 3 40

0.1

0.2

0.3

0.4

Period (sec)

At(

g)

lr = 3

lr = 5

0 1 2 3 40

2

4

6

8

10

12

14

Period (sec)

D (

in)

0 1 2 3 40

0.1

0.2

0.3

0.4

Period (sec)

At(

g)

lr = 3

lr = 5

190

(a) LA13, lr = 1/2 (b) LA13, lr = 1/4

(c) NF01, lr = 1/3 (d) NF01, lr = 1/5

(e) LS17c, lr = 1/3 (f) LS17c, lr = 1/5

Figure 4-21 Comparison of nonlinear spectra for Cyiso = 0.06, test data for FP type 1 and 3, and elastic 5% damped spectra, ground motions at peak span settings.

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

isolated system bridge deck model elastic, 5% damping

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

191

(a) LA13, lr = 1/2 (b) LA13, lr = 1/4

(c) NF01, lr = 1/3 (d) NF01, lr = 1/5

(e) LS17c, lr = 1/3 (f) LS17c, lr = 1/5

Figure 4-22 Comparison of nonlinear spectra for Cyiso = 0.06 and elastic 5% damped spectra, ground motions at peak span settings.

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

isolated system bridge deck model elastic, 5% damping

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

0 0.2 0.4 0.6 0.8 10

2

4

6

8

Tsub

(sec)

Dis

p. (

in)

192

4.12 Influence of Vertical Motions on Sliding Systems

The various configurations of the bridge deck specimen allowed the evaluation of the

influence of vertical motions on the response of a bridge specimen utilizing a sliding

isolation system. All configurations of the specimen incorporating FP bearings were tested

on the earthquake simulator with both uni-directional and bi-directional components of

the NF01_02 table motion with and without its vertical component.

Table 4-20 below presents complete results from these tests. For each of the test

sequences, peak deck, isolator, and substructure displacement and isolator force response

are shown for the same sequence performed with and without vertical input. The trend in

these results is illustrated in Table 4-21, where the ratio of response with vertical input to

the response without vertical input is tabulated for each sequence. A value greater than 1.0

indicates that the response is greater with the vertical signal, while a value less than 1.0

indicates the vertical signal decreased peak response. Mean statistics for the entire group

of tests are presented at the end of the table. The data indicates that the effect of vertical

input motion is small in general and tends to effect isolator and total displacement

response most significantly. Peak isolator and total displacement response were increased

by approximately 9 and 5 percent on average, respectively, by the addition of the vertical

input component. Peak substructure displacement and isolator shear force, on the other

hand, were effected by less than 3 and 1 percent on average, respectively. As noted earlier,

the FP bearing has a characteristic strength and stiffness dependant on the applied axial

force (i.e., Qd = µN and kd = N/R, see Equation 3, Chapter 14). As the axial load fluctuates

with the addition of the vertical signal, the bearing’s strength will vary more than its

stiffness (as 1/R is generally smaller than µ). The effect on deck and isolator

193

displacements (which are sensitive to system strength) would then be expected to be more

pronounced, as shear force response (and consequently substructure displacements and

isolator force response) was shown to be less sensitive to strength variations (see Section

2.3.5.2).

4.13 Effect of impact against bearing restraint system

Due to an error in setting the command signal, Configuration 1, test no. 5 (see Figure 4-

23), resulted FP bearing Q3 impacting against the restrainer ring. Due to the rigid nature

of the substructure in this case, the resulting impact produced a significant shear force

imposed locally on the foundations of this bearing. The effect of this impact on overall

shear force demand and displacement response was, however, less significant.

Figure 4-23 (a), (c), (e), and (g) show hysteretic response of each individual FP bearings in

this test. As seen in Figure 4-23 (e), the FP bearing at quadrant Q3 experienced a

significant increase in shear response as a result of the impact (an approximately 250

percent increase above the shear imposed immediately prior to the impact). In contrast,

Figure 4-23 (b) illustrates that total overall shear response increased less significantly as a

result (by approximately 30 percent). Figure 4-23 (f) illustrates the dramatic spike which

occurs during the shear time-history for the FP bearing at quadrant Q3. In comparison,

Figure 4-23 (h) shows the time-history response of global shear, illustrating how the

impact on total shear is less dramatic. Figure 4-23 (d) shows a time-history of

displacement response at the deck level of the bridge. This time-history illustrates how

insignificant the effect of this impact is on the global isolated response of the specimen.

194

Table 4-20 Response of isolated bridge deck on FP bearings with and without vertical motion input (nominal periods: To = 0 sec, T1 = 0.25 sec, T2 = 0.8 sec, T3 = 1 sec)

RecordTsubx-

Tsuby-

U w/o U w/ V w/o V w/ Uiso w/o Uiso w/ Usub w/o Usub w/

mpier/Mdeck =.05, FP type 3

NF01 T1 T1 4.315 4.276 15.785 16.012 4.136 4.144 0.224 0.206

NF02 T1 T1 3.357 3.706 12.672 12.233 3.284 3.656 0.091 0.101

NF01_02 T1 T1 5.363 5.377 17.299 17.586 5.231 5.248 0.176 0.182

NF01 T2 T3 5.471 5.560 15.387 15.259 4.175 4.280 1.360 1.392

NF02 T2 T3 6.009 6.258 15.472 15.447 3.553 4.004 2.528 2.512

NF01_02 T2 T3 7.945 8.565 19.284 19.280 5.360 5.972 2.875 2.891

mpier/Mdeck = 0.0, FP type 2

NF01 T0 T0 3.501 3.385 14.896 14.630

NF02 T0 T0 2.635 3.060 12.533 12.575

NF01_02 T0 T0 4.560 4.580 17.921 17.996


NF01 T2 T3 5.002 5.060 15.240 15.272 3.561 3.672 1.510 1.512

NF02 T2 T3 4.485 5.052 13.590 14.301 2.407 3.014 2.173 2.272

NF01_02 T2 T3 6.505 6.823 17.334 16.848 4.450 4.52 2.560 2.457


NF01 T1 T1 4.056 4.046 16.180 15.643 3.909 3.923 0.581 0.187

NF02 T1 T1 2.812 3.460 12.321 12.514 2.752= 3.394 0.079 0.092

NF01_02 T1 T1 5.035 5.0018 17.337 17.161 4.873 4.871 0.204 0.173

NF01 T2 T2 5.235 5.374 15.321 15.376 3.933 4.133 1.328 1.292

NF02 T2 T2 3.565 3.957 12.165 12.268 2.409 2.805 1.297 1.331

NF01_02 T2 T2 6.392 6.712 18.412 17.998 4.798 5.148 1.811 1.763

NF01 T2 T3 5.220 5.615 15.652 16.223 3.865 4.292 1.367 1.403

NF02 T2 T3 4.680 5.856 13.552 15.024 2.685 3.695 2.168 2.382

NF01_02 T2 T3 6.606 7.343 18.253 17.784 4.555 4.966 2.578 2.591


195

NF01 T2 T3 5.367 5.426 15.286 15.727 3.879 3.867 1.596 1.589

NF02 T2 T3 5.475 5.650 14.580 13.915 3.435 3.694 2.525 2.360

NF01_02 T2 T3 7.169 7.323 18.290 18.175 4.679 4.834 2.856 2.787


NF01_02S5 T0 T0 3.590 3.502 13.504 13.594

NF01_02S3 T0 T0 5.025 5.036 18.330 17.545


NF01_02S5 T0 T0 2.996 3.012 13.385 13.535

NF01_02S3 T0 T0 4.429 4.439 16.323 15.958

Table 4-21 Effect of vertical motion on peak response of isolated bridge deck on FP bearings (nominal periods: To = 0 sec, T1 = 0.25 sec, T2 = 0.8 sec, T3 = 1 sec)

RecordTsubx-

Tsuby-

U/Uo V/Vo Uiso/Uiso_0 Usub/Usub_0

mpier/Mdeck =.05, FP type 3

NF01 T1 T1 0.991 1.014 1.002 0.920

NF02 T1 T1 1.104 0.965 1.113 1.110

NF01_02 T1 T1 1.003 1.017 1.003 1.034

NF01 T2 T3 1.016 0.992 1.025 1.024

NF02 T2 T3 1.041 0.998 1.127 0.994

NF01_02 T2 T3 1.078 1.000 1.114 1.006


NF01 T0 T0 0.967 0.982

NF02 T0 T0 1.161 1.003

NF01_02 T0 T0 1.004 1.004


NF01 T2 T3 1.012 1.002 1.031 1.001

NF02 T2 T3 1.126 1.052 1.252 1.046

NF01_02 T2 T3 1.049 0.972 1.016 0.960

Table 4-20 Response of isolated bridge deck on FP bearings with and without vertical motion input (nominal periods: To = 0 sec, T1 = 0.25 sec, T2 = 0.8 sec, T3 = 1 sec)

196

The effect is minor presumably because the impact occurs during an instance of low

velocity in the bearing displacement cycle resulting in little change in energy.


NF01 T1 T1 0.998 0.967 1.004 0.322

NF02 T1 T1 1.230 1.016 1.233 1.165

NF01_02 T1 T1 0.993 0.990 1.000 0.848

NF01 T2 T2 1.027 1.004 1.051 0.973

NF02 T2 T2 1.110 1.008 1.164 1.026

NF01_02 T2 T2 1.050 0.978 1.073 0.973

NF01 T2 T3 1.076 1.036 1.110 1.026

NF02 T2 T3 1.251 1.109 1.376 1.099

NF01_02 T2 T3 1.112 0.974 1.090 1.005


NF01 T2 T3 1.011 1.029 0.997 0.996

NF02 T2 T3 1.032 0.954 1.075 0.935

NF01_02 T2 T3 1.021 0.994 1.033 0.976


NF01_02S5 T0 T0 0.975 1.007

NF01_02S3 T0 T0 1.002 0.957


NF01_02S5 T0 T0 1.005 1.011

NF01_02S3 T0 T0 1.002 0.978

Statistical scatter

Mean values 1.052 1.000 1.090 0.973

Standard deviation 0.072 0.032 .099 0.164

Table 4-21 Effect of vertical motion on peak response of isolated bridge deck on FP bearings (nominal periods: To = 0 sec, T1 = 0.25 sec, T2 = 0.8 sec, T3 = 1 sec)

197

This test sequence presents an opportunity to elaborate on the effect of impact on the

behavior of isolated bridge systems. As seen in this sequence, the effect of impact on the

bearing restrainer ring was to increase base shear demands (above that experienced by a

fully isolated system with no restraint). It follows that base shear demand would likely

increase (to a point) as the restraint limit was decreased, given the same earthquake

demand. Assuming the substructure to be designed to its elastic limit (µ=1) for the fully

isolated case, it could be postulated that substructure ductility demand may increase

dramatically as the velocity of the bearing increases upon impact (as would likely be the

case with decreasing restraint limits) as shear in the bearing drives displacement ductility

demand into the yielding substructure component. If substructure components do not yield

(as for systems with significant overstrength), designers would need to consider the effect

of increased local shear demands on connections and weak or brittle elements in the load

path. Thus, designing the substructure to yield at or before contact may be an effective

means of limiting these forces.

Displacement demands imposed on heavy flexible components by these local impacts

may also require special consideration. Simple idealizations which do not account for

higher mode contributions, may not detectthe occurrence of these local impacts. Damage

driven by these increased local force demands may cause premature local damage at

overall system displacements below that considered for behavior with no restraint limit.

More notably, earthquake demands representative of mean (DBE) seismic risk may

overlook the occurrence of impact for events above the mean (which have a 50 percent

likelihood of occurrence).

198

This test sequence points out some special considerations when impact against

displacement restraints occur. However, further analysis and testing is required to provide

a reliable basis for recommending improvements in bearing restraint design requirements.

4.14 Effect of bearing wear on systems employing FP bearings

Data was collected during the Configuration 1 through 9 test sequences to evaluate the

effect of bearing wear on system response. During each test sequence, one table motion

was selected as a basis and run as an initial test, at intervals during the sequence, and at the

conclusion of the sequence. Data collected from these repeated motion tests was evaluated

to determine changes in system response due to supposed “wear” in bearing materials.

Bearing and slider conditions were documented visually at the beginning and completion

of each test sequence to supplement the evaluation.

Table 4-22 shows peak specimen response through these sequences of bearing wear tests.

The number of time-history motions run between subsequent wear tests is denoted as

“History (n)”. Table 4-23 illustrates the effect of bearing wear by normalizing response by

initial test data. Results indicate that bearing wear effects overall system displacement

response only slightly, despite extreme wear in individual sliders. The effect of bearing

wear on shear response was similar, typically. However, an increase of nearly 20 percent

in force did occur during test sequence 9, albeit after 25 large test earthquake signals were

run. In some cases, more than 100 earthquakes were imposed without significant

deterioration in system performance.

199

(a) Q1 bearing hysteresis (b) Global bearing hysteresis

(c) Q2 bearing hysteresis (d) Global displacement history

(e) Q3 bearing hysteresis (f) Q3 shear history

Figure 4-23 Effect of impact against bearing restraint, FP type 1, Config. 1, Test No. 5, LS17c_18c, Span Setting [790/0/0]

−5 0 5−15

−10

−5

0

5

10

15

Disp. (in)

She

ar (

kips

)

−5 0 5−30

−20

−10

0

10

20

30

Disp. (in)

She

ar (

kips

)

−5 0 5−15

−10

−5

0

5

10

15

Disp. (in)

She

ar (

kips

)

4 6 8 10−8

−6

−4

−2

0

2

4

6

8

time (sec)

Dis

p. (

in)

−5 0 5−15

−10

−5

0

5

10

15

Disp. (in)

Sh

ea

r (k

ips)

4 4.5 5 5.5 6−15

−10

−5

0

5

10

15

time (sec)

Sh

ea

r (k

ips)

200

(g) Q4 bearing hysteresis (h) Global shear history

Figure 4-23 Effect of impact against bearing restraint, FP type 1, Config. 1, Test No. 5, LS17c_18c, Span Setting [790/0/0]

−5 0 5−15

−10

−5

0

5

10

15

Disp. (in)

She

ar (

kips

)

4 4.5 5 5.5 6−30

−20

−10

0

10

20

30

time (sec)

She

ar (

kips

)

201

Table 4-22 Peak system response after multiple signals

Sequence History (n) U V Uiso Usub

1

0 6.862 18.753

41 6.672 18.754

55 6.606 18.656

2

0 3.969 14.041

57 4.211 14.928

60 4.256 15.114

3

0 6.372 18.867

35 6.388 19.091

42 6.421 19.077

64 6.476 19.166

115 6.268 18.924

40 3.634 18.761

23 3.682 18.499

50 6.016 16.934 4.500 1.805

18 6.020 17.921 4.445 1.903

60 4.481 15.095 4.407 0.152

28 4.423 15.162 4.328 0.144

70 6.701 18.387 4.097 2.718

24 6.682 18.413 4.279 2.706

80 6.721 18.130 4.199 2.739

30 6.659 17.985 4.165 2.712

90 5.001 15.588 4.932 0.188

25 4.812 18.766 4.735 0.141

100 6.947 17.914 4.452 2.635

27 7.100 18.077 4.559 2.685

110 6.974 17.751 4.559 2.571

16 6.988 17.602 4.448 2.583

120 4.853 23.484

16 5.141 23.741

202

Table 4-23 Effect of bearing wear on peak system response

Sequence History (n) U/U0 V/V0 Uiso/Uiso_0 Usub/Usub

1

0 1.000 1.000

41 0.972 1.000

55 0.963 0.995

2

0 1.000 1.000

57 1.061 1.063

60 1.072 1.076

3

0 1.000 1.000

35 1.003 1.012

42 1.008 1.011

64 1.017 1.016

115 0.984 1.003

40 1.000 1.000

23 1.013 0.986

50 1.000 1.000 1.000 1.000

18 1.001 1.058 0.988 1.054

60 1.000 1.000 1.000 1.000

28 0.987 1.004 0.982 0.944

70 1.000 1.000 1.000 1.000

24 0.997 1.001 1.045 0.995

80 1.000 1.000 1.000 1.000

30 0.991 0.992 0.992 0.990

90 1.000 1.000 1.000 1.000

25 0.962 1.204 0.960 0.747

100 1.000 1.000 1.000 1.000

27 1.022 1.009 1.024 1.019

110 1.000 1.000 1.000 1.000

16 1.002 0.992 0.976 1.005

120 1.000 1.000

16 1.059 1.011

214

5 Displacement Estimates in Isolated Bridges

5.1 Introduction

The dynamic analyses carried out in this chapter are developed to refine and extend the

pilot analyses presented in Chapter 2 and the experimental studies presented in Chapters 3

and 4. This chapter focuses on the program objectives outlined in Section 1.3, specifically

program tasks 3, 5, and 7. The purpose of these tasks is to develop a more comprehensive

understanding of the behavior of simple isolated bridge overcrossings and to evaluate

current design procedures.

5.1.1 Problem Summary

This chapter was developed to systematically evaluate several design issues. Specifically

this chapter

• reviews the numerical models utilized for the subsequent analysis;

• describes the suites of ground motion time-histories used in the analyses;

• presents a systematic evaluation of the reliability of the Uniform Load Method;

• examines the adequacy of the provisions of the AASHTO Guide Specifications related to bi-directional ground motions;

• evaluates the ability of the Guide Specifications R-Factor design approach to control substructure yielding and the effect of this yielding on system response;

215

• describes a comprehensive parametric study of nonlinear isolated bridge response to identify trends related to structural and isolator properties and ground motion charac-teristics that can be used as guidelines to achieve improved performance; and

• evaluates the effect of local substructure mass and damping contributions to assess the adequacy of simplified single and multiple degree-of-freedom idealizations.

Detailed summaries of each analytical segment are included at the conclusion of each

section. Conclusions taken from these evaluations are utilized in Chapter 6 to formulate

overall recommendations for improvements of design procedures.

5.2 Analytical Bridge Models

Three simplified analytical models were utilized for the studies reported in this chapter as

summarized in Table 5-1.

The first two models were developed previously for the pilot analytical studies described

in Chapter 2 (see Section 2.3.3). Each of these models represents deck mass, isolation

system, and substructure properties as lumped components. The first model (Model 1) is a

planar two degree-of-freedom representation with dynamic degrees-of-freedom

representing deck and substructure displacements, respectively (see Figure 2-4(b)). This

model is similar to the simplified single degree-of-freedom model presented in Chapter 2

(see Figure 2-4(a)), but with the ability to incorporate distinct substructure mass and

damping contributions. The second model (Model 2) is a generalized multi-degree-of-

freedom, bi-directional bridge model with dynamic degrees-of-freedom representing deck

and substructure displacements in the x- & y-plane (see Figure 2-4(c)).

The final structural model (Model 3) is a planar single degree-of-freedom model, with the

contribution of a bilinear yielding hysteretic model for the substructure component

216

incorporated. Figure 5-1 presents a general description of this model. The model

represents the overall nonlinear behavior of an idealized isolated bridge responding uni-

directionally with yielding of the substructure and isolator explicitly represented.

Parameters are specified to match the component definitions of the AASHTO Guide

Specification with bilinear hysteretic isolation component behavior (see Chapter 2, Figure

2-2 (a)). Behavior of the substructure is also idealized as bilinear hysteretic, with a

specified yield strength and hardening stiffness ratio defined as shown (see Figure 5-1

(c)). Two internal system degrees of freedom represent the deck and substructure

component, respectively (see Figure 5-1 (b)). Substructure mass and damping

contributions are assumed to be negligible in this case such that overall system force-

deformation response is condensed to a single degree-of-freedom system with lumped

deck mass. This system has a tri-linear hysteresis determined from the mechanical

properties of the bilinear isolation and substructure hysteretic components acting in series

(see Figure 5-1 (d)).

5.3 Ground motion time-history suites

The nonlinear response of an isolated bridge overcrossing is influenced by many factors.

Foremost of these is perhaps the characteristics of the earthquake input. For purposes of

these analytical studies, several suites of ground motion time-history records were

compiled to facilitate evaluation of the effect of input characteristics on system response.

The features of each of these record suites were chosen considering the purpose of each

evaluation being undertaken. Five (5) different suites were utilized for the analytical

evaluations as summarized in Table 5-1. Some of these aggretized sets of records were

217

compiled to represent a uniform hazard of given design basis return interval at sites of a

particular standard soil class (i.e., design basis events representing the design spectra on

average). In other cases, de-aggretized sets of recorded ground motions were compiled to

represent discrete sets (bins) of magnitude and distance pairs for specific soil types. These

were selected to explicitly isolate the effect of earthquake magnitude, distance to fault

rupture and soil characteristics on response. In each case two horizontal components of the

record at each site were available in bi-directional ground motion time-history pairs. In

(a) Isolated bridge deck (b) Idealized dynamic model

(c) Substructure force-displacement (d) System force-displacement

Figure 5-1 Planar single-degree-of-freedom bridge model with yielding substructure

di

dsub

d

Yielding substructure bent

Isolation bearings

Mdeck F

Mdeck

d

kukd

Csub=0

dsub

di

KsubαKsub

Qd/(1-kd/ku)Fysub

bent isolator

mpier = 0

F=CsMg

Dsub

F

Disp.

CysubMg=Fysub

Ksub

αKsub

Fyiso

D

K1=kuKsub/(ku+Ksub)

K2=kdKsub/(kd+Ksub)

F

Disp.

K3= kdαKsubFysub(kd+αKsub)

218

some cases only one component was used, while in others both horizontal components

were considered.

Several synthetic impulsive records were also developed for these studies. These records

were created to represent the fundamental features of near-fault, shock pulses. Detailed

information on each of these ground motion suites is presented in the following sections.

5.3.1 Suites for evaluation of the AASHTO Uniform Load Method and R-factor provi-sions

For purposes of these evaluations, three suites of ground motion records were selected.

These suites were to developed to study the reliability of AASHTO Guide Spec provisions

related the the Uniform Load Method and R-factor approach under both far-field and near-

fault events.

The first suite is an aggretized set of records representing a uniform hazard for seismic

events in the Los Angeles area. This suite contains twenty ground motion records and

consists of ten, bi-directional, ground motion pairs. These records are representative of

Table 5-1 Summary of structural model and ground motion suitesused for analyses segments

Study Section Structural Model Ground Motion Suite

Uniform Load Method 5.4 Model 1Planar 2DOF

Uniform HazardSection 5.3.1

Bi-directional response 5.5 Model 2MDOF

Suite ASection 5.3.2

R-factor approach 5.6Model 3

2DOF w/ yielding substructure

Uniform HazardSection 5.3.1

Global parameter study 5.7 Model 1Planar 2DOF

Suite BSection 5.3.3

Local mass and damping effects 5.8 Model 1

Planar 2DOFSuite B

Section 5.3.3

219

stiff soil sites (NEHRP site type D). They have been amplitude and frequency scaled such

that their mean, 5% damped, elastic spectrum closely matches the USGS target spectrum

corresponding to a probability of exceedence of 10% in 50 years for the Los Angeles area

[SAC, 1997]. This probability of exceedance represents a return interval of 475 years,

approximately equal to the design basis event (i.e., DBE) targeted in the AASHTO

provisions for bridge designs [AASHTO, 1994]. Table 5-2 lists each of the individual

ground motions in this database suite, described as “LA10in50”.

Utilizing an acceleration coefficient of A = 0.4 (representative of a probability of

exceedence of 10 percent in 50 years in the San Fernando region [AASHTO, 1994]) and a

site soil coefficient of Si = 1.5675 (representing approximately AASHTO soil type II), the

ground motion records in this database were linearly scaled such that their mean spectrum

matched the AASHTO design spectrum for this set of parameters with minimum absolute

error, particularly in the constant velocity range (see Figure 5-2 (a)). This set of records

represents a suite of design basis earthquake motions compatible with the AASHTO

design spectrum (i.e., matching the design spectrum on average and the stipulated

probability of exceedance).

The second and third set of records compiled for this study are a set of ten, near-fault

recorded ground motion histories and a set of purely impulsive synthetic records. These

records represent fault-normal and fault-parallel, near-fault ground motion components.

Table 5-3 and 5-4 lists each of the individual ground motion records in the set of near-fault

and synthetic pulse records, described as “Near Fault” and “Impulsive”, respectively.

220

Figure 5-2(b) shows the mean spectra and statistical scatter in the “Near Fault” suite of

motions. Comparing this spectra with the LA10in50 mean spectra in Figure 5-2(a), it can

be seen that the “Near Fault” motions are richer in long period content on average,

suggestive of their more impulsive character. Figure 5-3 shows the ground acceleration,

velocity, and displacement time-histories for the first ten seconds of the NF09 earthquake

record from this suite. The impulsive content of this record is readily apparent,

particularly in the velocity and displacement histories. This characteristic is typical of the

ground motion records in the “Near-Fault” suite.

It has been suggested that near-fault earthquakes may be modeled with sufficient accuracy

(for structures responding at or near the peak in an earthquake’s response spectra) as

simple pulse motions [Krawinkler and Alavi, 1998]. On this basis, pure pulse motions

were developed for these studies with cosine and sine waves selected to model fault-

normal and fault-parallel displacement histories, respectively. Figure 5-2(c) shows the

spectra of this set of motions, titled “Impulsive”. Ground acceleration, velocity, and

displacement histories for these pulse records are shown in Figure 5-4. These records are

characteristic of pure pulse motions with an amplitude of A = 0.4 g and pulse duration of

Tp = 1.0 second. These values produced synthetic records with peak velocity and

displacement characteristics representative of actual near-fault earthquake records

[Krawinkler and Alavi, 1998]. The pulse duration of Tp = 1.0 second also represents the

mid-range of non-isolated bridge systems evaluated in this study (i.e., Tsub = 0 to 2

seconds), allowing an evaluation of isolated response for systems with periods shorter and

longer than the predominant pulse period.

221

Table 5-2 LA10in50 ground motion histories

Record ID Event Year Mag.1 Station Orientation2 R3 (km) Scale4

LA01 Imperial Valley 1940 6.9 El Centro, ivir 45 10 2.01LA02 Imperial Valley 1940 6.9 El Centro, ivir 135 10 2.01LA03 Imperial Valley 1979 6.5 El Centro, Array #5 45 4.1 1.01LA04 Imperial Valley 1979 6.5 El Centro, Array #5 135 4.1 1.01LA05 Imperial Valley 1979 6.5 El Centro, Array #6 45 1.2 0.84LA06 Imperial Valley 1979 6.5 El Centro, Array #6 135 1.2 0.84LA07 Landers 1992 7.3 Barstow 45 36 3.20LA08 Landers 1992 7.3 Barstow 135 36 3.20LA09 Landers 1992 7.3 Yermo 45 25 2.17LA10 Landers 1992 7.3 Yermo 135 25 2.17

LA11 Loma Prieta 1989 7.0 Gilroy, Array #3 45 12.4 1.79LA12 Loma Prieta 1989 7.0 Gilroy, Array #3 135 12.4 1.79LA13 Northridge 1994 6.7 Newhall 45 6.7 1.03LA14 Northridge 1994 6.7 Newhall 135 6.7 1.03LA15 Northridge 1994 6.7 Rinaldi RS 45 7.5 0.79LA16 Northridge 1994 6.7 Rinaldi RS 135 7.5 0.79LA17 Northridge 1994 6.7 Sylmar 45 6.4 0.99LA18 Northridge 1994 6.7 Sylmar 135 6.4 0.99

LA19 North Palm Springs 1986 6.0 dhsp 45 6.7 2.97

LA20 North Palm Springs 1986 6.0 dhsp 135 6.7 2.97

1. Moment magnitude

2. Orientation with respect to fault in degrees

3. Distance from fault

4. Amplitude scale applied to recorded ground motion history [SAC 1997]

222

Table 5-3 Near Fault ground motion histories

Record ID Event Year Mag.1 Station Orientation2 R3 (km)

Soil Type4

NF01 Tabas, Iran 1978 7.4 Tabas N 1.2 DNF02 Tabas, Iran 1978 7.4 Tabas P 1.2 DNF03 Loma Prieta 1989 7.0 Los Gatos N 3.5 DNF04 Loma Prieta 1989 7.0 Los Gatos P 3.5 DNF09 Erzincan, Turkey 1992 6.7 Erzincan N 2.0 DNF10 Erzincan, Turkey 1992 6.7 Erzincan P 2.0 DNF15 Northridge 1994 6.7 Olive View N 6.4 DNF16 Northridge 1994 6.7 Olive View P 6.4 DNF17 Kobe 1995 6.9 Kobe JMA N 3.4 DNF18 Kobe 1995 6.9 Kobe JMA P 3.4 D1. Moment magnitude

2. Orientation with respect to fault; N = normal; P = parallel


4. Soil classification per the 1997 NEHRP Recommended Provisions (FEMA 1997)

Table 5-4 Impulsive ground motion histories

Record ID Event Station Orientation1 Tp2

(sec)A3

(g)V4

(in/s)D5

(in)

FN Cosine Pulse Fault Normal 90 1.0 0.4 24.6 7.8FP Sine Pulse Fault Parallel 0 1.0 0.4 49.2 24.6

1. Orientation with respect to fault in degrees.

2. Pulse duration

3. Peak acceleration amplitude.

4. Peak velocity amplitude.

5. Peak displacement amplitude.

223

a. AASHTO design spectrum and mean spectra of LA10in50 database motions

b. Mean spectra of Near Fault database motions

c. Spectra of Impulsive database motions

Figure 5-2 Psuedo-acceleration spectra for suites of ground motion histories, 5% damping

0 1 2 3 40

1

2

3

4

Period (sec)

Psu

edo−

acce

lera

tion

(g) Mean

AASHTO

maximum

minimum

+−1σ

0 1 2 3 40

1

2

3

4

Period (sec)

Psu

edo−

acce

lera

tion

(g) Mean

maximum

minimum

+−1σ

0 1 2 3 40

0.5

1

1.5

Period (sec)

Psu

edo−

acce

lera

tion

(g) Cosine pulse

Sine pulse

224

a. Ground acceleration time-history

b. Ground velocity time-history

c. Ground displacement time-history

Figure 5-3 Time-histories for NF09 earthquake record (first 10 seconds)

0 1 2 3 4 5 6 7 8 9 10−0.5

0

0.5

Time (sec)

Acc

ele

ratio

n (

g)

0 1 2 3 4 5 6 7 8 9 10−50

0

50

Time (sec)

Vel

ocity

(in

/s)

0 1 2 3 4 5 6 7 8 9 10−20

−10

0

10

20

Time (sec)

Dis

plac

emen

t (in

)

225

a. Cosine pulse, acceleration time-history b. Sine pulse, acceleration time-historyc

c. Cosine pulse, velocity time-history d. Sine pulse, velocity time-history

e. Cosine pulse, displacement time-history f. Sine pulse, displacement time-history

Figure 5-4 Time-histories for Impulsive earthquake records

0 0.5 1 1.5 2−0.5

−0.25

0

0.25

0.5

Time (sec)

Acc

ele

ratio

n (

g)

0 0.5 1 1.5 2−50

−25

0

25

50

Time (sec)

Ve

loci

ty (

in/s

)

0 0.5 1 1.5 2−50

−25

0

25

50

Time (sec)

Ve

loci

ty (

in/s

)

0 0.5 1 1.5 2−30

−15

0

15

30

Time (sec)

Dis

pla

cem

en

t (in

)

0 0.5 1 1.5 2−30

−15

0

15

30

Time (sec)

Dis

pla

cem

en

t (in

)

226

5.3.2 Suites for evaluation of AASHTO provisions related to bi-directional response

Fifty bi-directional pairs of ground motion time-histories, entitled Suite A were selected

for this study. These one-hundred motions were classified into five bins of twenty motions

each (i.e., ten pairs per bin) grouped by magnitude, distance to active fault, and soil type.

This classification provides a de-aggretization of these effects on the ground motion

characteristics. These bin classifications are outlined in Table 5-5. Table 5-6 through Table

5-10 list each of the individual ground motions in Bin 1 through 5, respectively. Bin 1

motions were selected from the ground motion database developed for the SAC Joint

Venture project [SAC, 1997]. Bin 2 through 5 motions were selected from the PEER

Strong Motion Database [PEER, 2000].

For each earthquake event and recording station, a pair of bi-directional ground motion

time histories are listed. For Bin 1, the first (i.e., x-direction) and second (i.e., y-direction)

orthogonal components are oriented normal and parallel, respectively, to the active

earthquake fault. For Bin 2 through 5, the orientation of the recorded bi-directional

components are based upon the orientation of the recording instrument and represent a

somewhat random directivity to the active fault.

Figure 5-5 shows the mean psuedo-acceleration spectrum computed for the set of twenty

histories for each of the Bin 1 through 5 ground motion suites (for a damping ratio of 5%).

Maxima, minima, and +1σ statistics are shown on each plot to illustrate the distribution in

the sets of ground motion data. Note the increase in average amplitude and preponderance

of long period content for motions of increasing magnitude and/or decreasing distance to

the active fault.

227

Figure 5-6 (a) and (b) present the AASHTO psuedo-acceleration and displacement design

spectra, respectively, for selected values of A = 0.4 and Si = 1.5 (parameters representative

of a design basis probability of exceedence of 10% in 50 years in the San Fernando region

and AASHTO soil type II, respectively [AASHTO, 1999]). Figure 5-6 (c) shows a

comparison of the mean spectra of bi-directional bin motions to the spectral shape

represented by the AASHTO psuedo-acceleration design spectrum for three selected

values of A and Si. It is seen in this figure that the mean spectra represented by the Bin 1

through 5 motions fit well to the AASHTO spectral shape, particularly on the descending

branch, or velocity sensitive region. This similarity is important. The AASHTO design

procedures are postulated on the mean response of systems to motions compatible with

this spectral shape (i.e., motions which represent the design basis hazard level and

“match” the spectrum closely on average). It follows that mean response characteristics

for each of these motion bins may be interpreted to apply generally to the design

procedures in the AASHTO Guide Specification.

Figure 5-7 shows a comparison for this suite of records of the mean pseudo-acceleration

spectrum computed for all ground motion histories in a bin to the mean spectra computed

for the first (x-direction) and second (y-direction) component histories in the same bin.

Figure 5-7 (a) shows these results for Bin 1 motions, where the first- and second-

component histories represent fault-normal and fault-parallel components, respectively.

For this plot the effect of directivity on the spectra is readily apparent. This bin, which

represents larger magnitude earthquakes for sites nearer the active fault, the difference

between the “larger”, or fault normal, component and the mean spectrum representing

“average” directivity is very pronounced. This is consistent with trends of near-fault

228

motions discussed in earlier work by Somerville [Somerville, 1997]. Analyses performed

utilizing the Bin 1 motions may then be interpreted as accounting for the effect of fault

directivity explicitly. This same comparison is shown in Figure 5-7 (b) through (e) for Bin

2 through 5 motions. For these bins, bi-directional ground motion pairs are oriented with

random directivity, and the difference between the first- and second-component mean

spectra and the mean spectrum of all motions in the bin is nearly negligible, as expected.

For these ground motion histories, therefore, results will be interpreted to apply generally

to the spectrum representing “average” or random directivity only.

Table 5-5 Ground motion bin classification for Suite A

BIN Name Magnitude R(km) Soil Type Classification1 NF 6.7 - 7.4 < 10 D NEHRP2 LMSR 6.7 - 7.3 10 - 30 A,C USGS3 LMLR 6.7 - 7.3 30 - 60 A,C USGS4 SMSR 5.8 - 6.5 10 - 30 A,C USGS5 SMLR 5.8 - 6.5 30 - 60 A,C USGS

229

Table 5-6 Suite A, Bin 1 ground motion histories: near fault (NF)


Soil4Type

NF01 Tabas, Iran 1978 7.4 Tabas N 1.2 DNF02 Tabas, Iran 1978 7.4 Tabas P 1.2 DNF03 Loma Prieta 1989 7.0 Los Gatos N 3.5 DNF04 Loma Prieta 1989 7.0 Los Gatos P 3.5 DNF05 Loma Prieta 1989 7.0 Lex. Dam N 6.3 DNF06 Loma Prieta 1989 7.0 Lex. Dam P 6.3 DNF07 Cape Mendocino 1992 7.1 Petrolia N 8.5 DNF08 Cape Mendocino 1992 7.1 Petrolia P 8.5 DNF09 Erzincan, Turkey 1992 6.7 Erzincan N 2.0 DNF10 Erzincan, Turkey 1992 6.7 Erzincan P 2.0 D

NF11 Landers 1992 7.3 Lucerne N 1.1 DNF12 Landers 1992 7.3 Lucerne P 1.1 DNF13 Northridge 1994 6.7 Rinaldi N 7.5 DNF14 Northridge 1994 6.7 Rinaldi P 7.5 DNF15 Northridge 1994 6.7 Olive View N 6.4 DNF16 Northridge 1994 6.7 Olive View P 6.4 DNF17 Kobe 1995 6.9 Kobe JMA N 3.4 DNF18 Kobe 1995 6.9 Kobe JMA P 3.4 DNF19 Kobe 1995 6.9 Takatori N 4.3 DNF20 Kobe 1995 6.9 Takatori P 4.3 D1. Moment magnitude




230

Table 5-7 Suite A, Bin 2 ground motion histories: large magnitude small distance (LMSR)

Record ID Event Year Mag.1 Station Orientation2 R3 (km) Soil4Type

LP89go1x Loma Prieta 1989 6.9 Gilroy Array #1 0 11.2 A

LP89go1y Loma Prieta 1989 6.9 Gilroy Array #1 90 11.2 A

LP89sgix Loma Prieta 1989 6.9 Hollister: SAGO Vault 270 30.6 A

LP89sgiy Loma Prieta 1989 6.9 Hollister: SAGO Vault 360 30.6 A

NR94lo9x Northridge 1994 6.7 Lake Hughes #9 0 26.8 A

NR94lo9y Northridge 1994 6.7 Lake Hughes #9 90 26.8 A

NR94wonx Northridge 1994 6.7 LA: Wonderland Ave. 95 22.7 A

NR94wony Northridge 1994 6.7 LA: Wonderland Ave. 185 22.7 A

SF71lo9x San Fernando 1971 6.6 Lake Hughes #9 21 23.5 A

SF71lo9y San Fernando 1971 6.6 Lake Hughes #9 291 23.5 A

LP89go2x Loma Prieta 1989 6.9 Gilroy Array #2 0 12.7 C

LP89go2y Loma Prieta 1989 6.9 Gilroy Array #2 90 12.7 C

LN92yerx Landers 1992 7.3 Yermo Fire Station 270 24.9 C

LN92yery Landers 1992 7.3 Yermo Fire Station 360 24.9 C

KB95abnx Kobe 1995 6.9 Abeno 0 23.8 C

KB95abny Kobe 1995 6.9 Abeno 90 23.8 C

IV79he1x Imperial Valley 1979 6.5 El Centro Array #1 140 15.5 C

IV79he1y Imperial Valley 1979 6.5 El Centro Array #1 230 15.5 C

NR94cnpx Northridge 1994 6.7 Canoga Pk: Topanga Can. 106 15.8 C

NR94cnpy Northridge 1994 6.7 Canoga Pk: Topanga Can. 196 15.8 C1. Moment magnitude



4. Soil classification per USGS

231

Table 5-8 Suite A, Bin 3 ground motion histories: large magnitude large distance (LMLR)


KB95chyx Kobe 1995 6.9 o Chihaya 0 48.7 A

KB95chyy Kobe 1995 6.9 o Chihaya 90 48.7 A

LN9229px Landers 1992 7.3 Twentynine Palms 0 42.2 A

LN9229py Landers 1992 7.3 Twentynine Palms 90 42.2 A

LP89mchx Loma Prieta 1989 6.9 Monterey City Hall 0 44.8 A

LP89mchy Loma Prieta 1989 6.9 Monterey City Hall 90 44.8 A

NR94mtwx Northridge 1994 6.7 Mt. Wilson: CIT Seis Sta. 0 36.1 A

NR94mtwy Northridge 1994 6.7 Mt. Wilson: CIT Seis Sta. 90 36.1 A

NR94grnx Northridge 1994 6.7 San Gabriel: E. Grand Ave. 180 41.7 A

NR94grny Northridge 1994 6.7 San Gabriel: E. Grand Ave. 270 41.7 A

KB95tdox Kobe 1995 6.9 o Tadoka 0 30.5 C

KB95tdoy Kobe 1995 6.9 o Tadoka 90 30.5 C

LN92psax Landers 1992 7.3 Palm Springs Airport 0 37.5 C

LN92psay Landers 1992 7.3 Palm Springs Airport 90 37.5 C

LP89slcx Loma Prieta 1989 6.9 Palo Alto: SLAC Lab 270 36.3 C

LP89slcy Loma Prieta 1989 6.9 Palo Alto: SLAC Lab 360 36.3 C

NR94casx Northridge 1994 6.7 Compton: Castlegate Ave. 0 49.6 C

NR94casy Northridge 1994 6.7 Compton: Castlegate Ave. 90 49.6 C

IV79vctx Imperial Valley 1979 6.5 Victoria 75 54.1 C

IV79vcty Imperial Valley 1979 6.5 Victoria 345 54.1 C1. Moment magnitude




232

Table 5-9 Suite A, Bin 4 ground motion histories: small magnitude small distance (SMSR)


MH84go1x Morgan Hill 1984 6.2 Gilroy Array #1 230 16.2 A

MH84go1y Morgan Hill 1984 6.2 Gilroy Array #1 320 16.2 A

PS86silx North Palm Springs 1986 6.0 Silent Valley: Poppet F 0 25.8 A

PS86sily North Palm Springs 1986 6.0 Silent Valley: Poppet F 90 25.8 A

WH87wonx Whittier Narrows 1987 6.0 LA: Wonderland

Ave. 75 24.6 A

WH87wony Whittier Narrows 1987 6.0 LA: Wonderland

Ave. 165 24.6 A

WH87mtwx Whittier Narrows 1987 6.0 Mt. Wilson: CIT Seis. Sta. 0 21.2 A

WH87mtwy Whittier Narrows 1987 6.0 Mt. Wilson: CIT Seis. Sta. 90 21.2 A

CL79go1x Coyote Lake 1979 5.7 Gilroy Array #1 230 9.3 A

CL79go1y Coyote Lake 1979 5.7 Gilroy Array #1 320 9.3 A

CL79hvrx Coyote Lake 1979 5.7 Halls Valley 150 31.2 C

CL79hvry Coyote Lake 1979 5.7 Halls Valley 240 31.2 C

IV79cxox Imperial Valley 1979 6.5 Calexico Fire Station 225 10.6 C

IV79cxoy Imperial Valley 1979 6.5 Calexico Fire Station 315 10.6 C

MH84go2x Morgan Hill 1984 6.2 Gilroy Array #2 0 15.1 C

MH84go2y Morgan Hill 1984 6.2 Gilroy Array #2 90 15.1 C

LM80srmx Livermore 1980 5.8 San Ramon Fire Station 70 21.7 C

LM80srmy Livermore 1980 5.8 San Ramon Fire Station 340 21.7 C

WH87buex Whittier Narrows 1987 6.0 Burbank: N. Buena Vista 250 23.7 C

WH87buey Whittier Narrows 1987 6.0 Burbank: N. Buena Vista 340 23.7 C1. Moment magnitude




233

Table 5-10 Suite A, Bin 5 ground motion histories: small magnitude large distance (SMLR)


PS86azfx North Palm Springs 1986 6.0 Anza Fire Station 225 46.7 A

PS86azfy North Palm Springs 1986 6.0 Anza Fire Station 315 46.7 A

PS86armx North Palm Springs 1986 6.0 Anza: Red Mountain 270 45.6 A

PS86army North Palm Springs 1986 6.0 Anza: Red Mountain 360 45.6 A

PS86ho2x North Palm Springs 1986 6.0 Winchester Bergman Ran 0 57.6 A

PS86ho2y North Palm Springs 1986 6.0 Winchester Bergman Ran 90 57.6 A

PS86ho1x North Palm Springs 1986 6.0 Murnetz Hot Springs 0 63.3 A

PS86ho1y North Palm Springs 1986 6.0 Murnetz Hot Springs 90 63.3 A

SN71sodx San Fernando 1971 6.6 Upland: San Antonio Dam 15 58.1 B5

SN71sody San Fernando 1971 6.6 Upland: San Antonio Dam 285 58.1 B5

CA83co8x Coalinga 1983 6.4 Parkfield: Cholame 8k1 0 50.7 C

CA83co8y Coalinga 1983 6.4 Parkfield: Cholame 8k1 270 50.7 C

PS86ho6x North Palm Springs 1986 6.0 San Jacinto Valley Cem. 270 39.6 C

PS86ho6y North Palm Springs 1986 6.0 San Jacinto Valley Cem. 360 39.6 C

MH84hchx Morgan Hill 1984 6.2 Hollister City Hall 1 32.5 C

MH84hchy Morgan Hill 1984 6.2 Hollister City Hall 271 32.5 C

WN87cnpx Whittier Narrows 1987 6.0 Canoga Park: Topanga Can. 106 47.4 C

WN87cnpy Whittier Narrows 1987 6.0 Canoga Park: Topanga Can. 196 47.4 C

LM80stpx Livermore 1980 5.8 Tracy: Sew. Treat. Plant 93 37.3 C

LM80stpy Livermore 1980 5.8 Tracy: Sew. Treat. Plant 183 37.3 C1. Moment magnitude




5. Soil classification Type A per Geomatrix

234

a. Bin 1 - NF

b. Bin 2- LMSR

c. Bin 3 - LMLR

Figure 5-5 Mean psuedo-acceleration spectra: Suite A, 5% damping

0 0.5 1 1.5 2 2.5 3 3.5 40

1

2

3

4

Period (sec)

Psu

edo−

acce

lera

tion

(g) mean

max/min

+-1σ, typical

0 0.5 1 1.5 2 2.5 3 3.5 4−0.5

0

0.5

1

1.5

2

2.5

Period (sec)

Psu

edo−

acce

lera

tion

(g) mean

max/min

0 0.5 1 1.5 2 2.5 3 3.5 4−0.5

0

0.5

1

1.5

Period (sec)

Psu

edo−

acce

lera

tion

(g) mean

max/min

235

d. Bin 4 - SMSR

e. Bin 5 - SMLR

Figure 5-5 Mean psuedo-acceleration spectra: Suite A, 5% damping

0 0.5 1 1.5 2 2.5 3 3.5 4−0.5

0

0.5

1

Period (sec)

Psu

edo−

acce

lera

tion

(g) mean

max/min

0 0.5 1 1.5 2 2.5 3 3.5 4−0.2

0

0.2

0.4

0.6

Period (sec)

Psu

edo−

acce

lera

tion

(g) mean

max/min

236

a. AASHTO psuedo-acceleration spectrum b. AASHTO displacement spectrum

c. Mean spectra vs. AASHTO spectral shape

Figure 5-6 AASHTO elastic response spectra with comparison to mean psuedo-accel-eration spectra for Suite A ground motions, 5% damping

0 1 2 3 40

0.2

0.4

0.6

0.8

1

1.2

Period (sec)

Psu

ed

o−

acc

ele

ratio

n (

g)

Ag = 0.4gSi = 1.5, Soil Type II

varies as 1/T

0 1 2 3 40

5

10

15

20

25

Period (sec)

Dis

pla

cem

en

t (in

)


0 0.5 1 1.5 2 2.5 3 3.5 40

0.3

0.6

0.9

1.2

1.5

1.8

Period (sec)

Psu

edo−

acce

lera

tion

(g) AASHTO

BIN motions

varies as 1/T

237

a. Bin 1 - NF

b. Bin 2- LMSR

c. Bin 3 - LMLRFigure 5-7 Mean psuedo-acceleration spectra of Suite A ground motion bins vs. mean of first (x-direction) and second (y-direction) component histories, 5% damping

0 0.5 1 1.5 2 2.5 3 3.5 40

0.5

1

1.5

2

Period (sec)

Psu

edo−

acce

lera

tion

(g) mean of all Bin motions

mean of fault normal components mean of fault parallel components

0 0.5 1 1.5 2 2.5 3 3.5 40

0.2

0.4

0.6

0.8

Period (sec)

Psu

edo−

acce

lera

tion


mean of first components mean of second components

0 0.5 1 1.5 2 2.5 3 3.5 40

0.1

0.2

0.3

0.4

0.5

Period (sec)

Psu

edo−

acce

lera

tion



238

d. Bin 4 - SMSR

e. Bin 5 - SMLRFigure 5-7 Mean psuedo-acceleration spectra of Suite A ground motion bins vs. mean of first (x-direction) and second (y-direction) component histories, 5% damping

0 0.5 1 1.5 2 2.5 3 3.5 40

0.1

0.2

0.3

0.4

Period (sec)

Psu

edo−

acce

lera

tion



0 0.5 1 1.5 2 2.5 3 3.5 40

0.05

0.1

0.15

0.2

0.25

Period (sec)

Psu

edo−

acce

lera

tion



239

5.3.3 Suites for parametric studies of isolated bridge response and influence of substruc-ture mass and damping

One hundred ground motion time histories, entitled Suite B, were selected for these

studies. These motions were classified into five de-aggretized bins of twenty motions

each. Bins were classified by earthquake magnitude and distance to active fault as

described in Table 5-11. Bin 1 motions for this suite were taken from the ground motion

database developed for the SAC Joint Venture project [SAC, 1997] while Bin 2 through 5

motions were selected from the PEER Strong Motion Database [PEER, 2000].

Table 5-12 through Table 5-16 lists each of the individual ground motions in Bin 1 through

5, respectively. For Bin 1, ground motions are listed sequentially in bi-directional pairs

where the first and second orthogonal components are oriented normal and parallel to the

active fault, respectively. For Bin 2 through 5, only uni-directional components were

selected representing somewhat random fault directivity.

For these studies computational time constraints dictated that only half of the compiled

ground motion records could be used. To facilitate this, every other ground motion record

from each bin was selected for the parametric evaluations (beginning with the first record

of the bin). This implies that for Bin 1 motions, only the “stronger” fault-normal motions

were utilized these studies. Figure 5-8 shows the mean pseudo-acceleration spectrum

computed for the set of ten histories utilized from the Bin 1 through 5 ground motion

suites (for a damping ratio of 5%). Maxima, minima, and +1σ statistics are shown on each

plot to illustrate the distribution in the sets of ground motion data. Figure 5-9 presents a

comparison of these mean spectra. Note the increase in average amplitude and

preponderance of long period content as ground motion magnitude increases and/or

240

distance to the active fault decreases as for the previous de-aggretized, bi-directional bins

(see Figure 5-5).

As for the Suite A motions, Figure 5-10 compares the AASHTO design spectral shape to

the mean spectra of the Suite B motions. It is seen in this figure that the mean spectra

represented by these bins also fit well to the AASHTO spectral shape. This implies, as for

the Suite A motions, that mean response characteristics for each of these motion bins may

be interpreted to apply generally to the AASHTO Guide Spec procedures (see Section 5.3.2

above).

Table 5-11 Ground motion bin classification for Suite B

BIN Name Magnitude R(km) Soil Type Classification1 NF 6.7 - 7.4 < 10 D NEHRP2 LMSR 6.7 - 7.3 10 - 30 D NEHRP3 LMLR 6.7 - 7.3 30 - 60 D NEHRP4 SMSR 5.8 - 6.5 10 - 30 D NEHRP5 SMLR 5.8 - 6.5 30 - 60 D NEHRP

241

Table 5-12 Suite B, Bin 1 ground motion histories: near fault (NF)


Soil4Type

NF01 Tabas, Iran 1978 7.4 Tabas N 1.2 DNF02 Tabas, Iran 1978 7.4 Tabas P 1.2 DNF03 Loma Prieta 1989 7.0 Los Gatos N 3.5 DNF04 Loma Prieta 1989 7.0 Los Gatos P 3.5 DNF05 Loma Prieta 1989 7.0 Lex. Dam N 6.3 DNF06 Loma Prieta 1989 7.0 Lex. Dam P 6.3 DNF07 Cape Mendocino 1992 7.1 Petrolia N 8.5 DNF08 Cape Mendocino 1992 7.1 Petrolia P 8.5 DNF09 Erzincan, Turkey 1992 6.7 Erzincan N 2.0 DNF10 Erzincan, Turkey 1992 6.7 Erzincan P 2.0 D

NF11 Landers 1992 7.3 Lucerne N 1.1 DNF12 Landers 1992 7.3 Lucerne P 1.1 DNF13 Northridge 1994 6.7 Rinaldi N 7.5 DNF14 Northridge 1994 6.7 Rinaldi P 7.5 DNF15 Northridge 1994 6.7 Olive View N 6.4 DNF16 Northridge 1994 6.7 Olive View P 6.4 DNF17 Kobe 1995 6.9 Kobe JMA N 3.4 DNF18 Kobe 1995 6.9 Kobe JMA P 3.4 DNF19 Kobe 1995 6.9 Takatori N 4.3 DNF20 Kobe 1995 6.9 Takatori P 4.3 D1. Moment magnitude




242

Table 5-13 Suite B, Bin 2 ground motion histories: large magnitude small distance (LMSR)

Record ID Event Year Mag.1 Station Mechanism2 R3 (km) Soil4Type

IV40elc Imperial Valley 1940 7.0 El Centro Array #9 strike-slip 12.0 DLD92yer Landers 1992 7.3 Yermo Fire Station strike-slip 24.9 D

LP89agw Loma Prieta 1989 6.9 Agnews State Hospital reverse-oblique 28.2 D

LP89cap Loma Prieta 1989 6.9 Capitola reverse-oblique 14.5 D

LP89g03 Loma Prieta 1989 6.9 Gilroy Array #3 reverse-oblique 14.4 D

LP89g04 Loma Prieta 1989 6.9 Gilroy Array #4 reverse-oblique 16.1 D

LP89gmr Loma Prieta 1989 6.9 Gilroy Array #7 reverse-oblique 24.2 D

LP89hch Loma Prieta 1989 6.9 Hollister City Hall reverse-oblique 28.2 D

LP89hda Loma Prieta 1989 6.9 Hollister Diff. Array reverse-oblique 25.8 D

LP89svl Loma Prieta 1989 6.9 Sunnyvale: Colton Ave reverse-oblique 28.8 D

NR94cnp Northridge 1994 6.7 Canoga Park: Topanga Can reverse-slip 15.8 D

NR94far Northridge 1994 6.7 LA - N Faring Rd reverse-slip 23.9 DNR94fle Northridge 1994 6.7 LA - Fletcher Dr reverse-slip 29.5 DNR94glp Northridge 1994 6.7 Glendale - Las Palmas reverse-slip 25.4 DNR94hol Northridge 1994 6.7 Hollywood Store FF reverse-slip 25.5 DNR94stc Northridge 1994 6.7 17645 Saticoy St reverse-slip 13.3 DSF71pel San Fernando 1971 6.6 LA: Hollywood Store Lot reverse-slip 21.2 DSH87bra Superstition Hills 1987 6.7 Brawley strike-slip 18.2 DSH87icc Superstition Hills 1987 6.7 El Centro Imp. Co. Cent strike-slip 13.9 D

SH87wsm Superstition Hills 1987 6.7 Westmorland Fire Station strike-slip 13.3 D1. Moment magnitude

2. Faulting mechanism



243

Table 5-14 Suite B, Bin 3 ground motion histories: large magnitude large distance (LMLR)


BM68elc Borrego Mountain 1968 6.8 El Centro Array #9 strike-slip 46.0 DLD92ind Landers 1992 7.3 Indio: Coachella Canal strike-slip 55.7 DLD92psa Landers 1992 7.3 Palm Springs Airport strike-slip 37.5 D

LP89a2e Loma Prieta 1989 6.9 Hayward Muir Schoolreverse-oblique 57.4 D

LP89fms Loma Prieta 1989 6.9 Freemont: Emerson Ct

reverse-oblique 42.4 D

LP89hvr Loma Prieta 1989 6.9 Halls Valley reverse-oblique 31.6 D

LP89sjw Loma Prieta 1989 6.9 Salinas: John & Work reverse-oblique 32.6 D

LP89slc Loma Prieta 1989 6.9 Palo Alto: SLAC Lab reverse-oblique 36.3 D

NR94ana Northridge 1994 6.7 Anaverde Valley: City R reverse-slip 38.4 DNR94bad Northridge 1994 6.7 Covina: W. Badillo reverse-slip 56.1 DNR94cas Northridge 1994 6.7 Compton: Castlegate St reverse-slip 49.6 DNR94cen Northridge 1994 6.7 LA: Centinela St reverse-slip 30.9 DNR94cmr Northridge 1994 6.7 Camarillo reverse-slip 36.5 DNR94del Northridge 1994 6.7 Lakewood: Del Amo Blvd reverse-slip 59.3 D

NR94dwn Northridge 1994 6.7 Downey: Co Maint Bldg reverse-slip 47.6 DNR94eli Northridge 1994 6.7 Elizabeth Lake reverse-slip 37.2 DNR94jab Northridge 1994 6.7 Bell Gardens: Jaboneria reverse-slip 46.6 DNR94lh1 Northridge 1994 6.7 Lake Hughes #1 reverse-slip 36.3 DNR94loa Northridge 1994 6.7 Lawndale: Osage Ave reverse-slip 42.4 DNR94lv2 Northridge 1994 6.7 Leona Valley #2 # reverse-slip 37.7 D

1. Moment magnitude




244

Table 5-15 Suite B, Bin 4 ground motion histories: small magnitude small distance (SMSR)


IV79cal Imperial Valley 1979 6.5 Calipatria Fire Station strike-slip 23.8 DIV79chi Imperial Valley 1979 6.5 Chihuahua strike-slip 28.7 DIV79e01 Imperial Valley 1979 6.5 El Centro Array #1 strike-slip 15.5 DIV79e12 Imperial Valley 1979 6.5 El Centro Array #12 strike-slip 18.2 DIV79e13 Imperial Valley 1979 6.5 El Centro Array #13 strike-slip 21.9 DIV79qkp Imperial Valley 1979 6.5 Cucapah strike-slip 23.6 DIV79wsm Imperial Valley 1979 6.5 Westmorland Fire Station strike-slip 15.1 DLV80kod Livermore 1980 5.8 San Ramon Fire Station strike-slip 21.7 D

LV80srm Livermore 1980 5.8 San Ramon: Eastman Kodak strike-slip 17.6 D

MH84agw Morgan Hill 1984 6.2 Agnews State Hospital strike-slip 29.4 DMH84g02 Morgan Hill 1984 6.2 Gilroy Array #2 strike-slip 15.1 DMH84g03 Morgan Hill 1984 6.2 Gilroy Array #3 strike-slip 14.6 DMH84gmr Morgan Hill 1984 6.2 Gilroy Array #7 strike-slip 14.0 DPM73phn Point Mugu 1973 5.8 Port Hueneme reverse-slip 25.0 DPS86psa N. Palm Springs 1986 6.0 Palm Springs Airport strike-slip 16.6 D

WN87cas Whittier Narrows 1987 6.0 Compton: Castlegate St. reverse 16.9 DWN87cat Whittier Narrows 1987 6.0 Carson: Catskill Ave. reverse 28.1 DWN87flo Whittier Narrows 1987 6.0 Brea: S Flower Ave. reverse 17.9 D

WN87w70 Whittier Narrows 1987 6.0 LA: W 70th St. reverse 16.3 DWN87wat Whittier Narrows 1987 6.0 Carson: Water St. reverse 24.5 D

1. Moment magnitude




245

Table 5-16 Suite B, Bin 5 ground motion histories: small magnitude large distance (SMLR)


BO42elc Borrego 1942 6.5 El Centro Array #9 not given 49.0 D

CO83c05 Coalinga 1983 6.4 Parkfield: Cholame 5Wreverse-oblique 47.3 D

CO83c08 Coalinga 1983 6.4 Parkfield: Cholame 8Wreverse-oblique 50.7 D

IV79cc4 Imperial Valley 1979 6.5 Coachella Canal #4 strike-slip 49.3 DIV79cmp Imperial Valley 1979 6.5 Compuertas strike-slip 32.6 DIV79dlt Imperial Valley 1979 6.5 Delta strike-slip 43.6 DIV79nil Imperial Valley 1979 6.5 Niland Fire Station strike-slip 35.9 DIV79pls Imperial Valley 1979 6.5 Plaster City strike-slip 31.7 DIV79vct Imperial Valley 1979 6.5 Victoria strike-slip 54.1 D

LV80stp Livermore 1980 5.8 Tracy: Sewage Treat. Plant strike-slip 37.3 D

MH84cap Morgan Hill 1984 6.2 Capitola strike-slip 38.1 DMH84hch Morgan Hill 1984 6.2 Hollister City Hall strike-slip 32.5 DMH84sjb Morgan Hill 1984 6.2 San Juan Bautista strike-slip 30.3 C

PS86h06 N. Palm Springs 1986 6.0 San Jacinto Valley Ceme-tery strike-slip 39.6 D

PS86ino N. Palm Springs 1986 6.0 Indio strike-slip 39.6 D

WN87bir Whittier Narrows 1987 6.0 Downey - Birchdale # reverse 56.8 D

WN87cts Whittier Narrows 1987 6.0 LA: Century City CC South reverse 31.3 D

WN87har Whittier Narrows 1987 6.0 LB - Harbor Admin FF reverse 34.2 D

WN87sse Whittier Narrows 1987 6.0 Terminal Island: S Seaside reverse 35.7 D

WN87stc Whittier Narrows 1987 6.0 Northridge: Saticoy St. reverse 39.8 D

1. Moment magnitude




246

a. Bin 1 - NF (mean spectra of ten selected motions)

b. Bin 2- LMSR (mean spectra of ten selected motions)

c. Bin 3 - LMLR (mean spectra of ten selected motions)

Figure 5-8 Mean pseudo-acceleration spectra: Suite B, 5% damping

0 0.5 1 1.5 2 2.5 3 3.5 4−1

0

1

2

3

4

Period (sec)

Psu

edo−

acce

lera

tion

(g) mean

max/min

+-1σ, typical

0 0.5 1 1.5 2 2.5 3 3.5 4−0.5

0

0.5

1

1.5

Period (sec)

Psu

edo−

acce

lera

tion

(g) mean

max/min

0 0.5 1 1.5 2 2.5 3 3.5 4−0.2

0

0.2

0.4

0.6

Period (sec)

Psu

edo−

acce

lera

tion

(g) mean

max/min

247

d. Bin 4 - SMSR (mean spectra of ten selected motions)

e. Bin 5 - SMLR (mean spectra of ten selected motions)

Figure 5-8 Mean pseudo-acceleration spectra: Suite B, 5% damping

0 0.5 1 1.5 2 2.5 3 3.5 4−0.5

0

0.5

1

1.5

2

Period (sec)

Psu

edo−

acce

lera

tion

(g) mean

max/min

0 0.5 1 1.5 2 2.5 3 3.5 4−0.2

0

0.2

0.4

0.6

0.8

Period (sec)

Psu

edo−

acce

lera

tion

(g) mean

max/min

248

a. Bin 1, 2, and 3 (mean spectra of ten selected motions)

b. Bin 2, 3, 4, and 5 (mean spectra of ten selected motions)

Figure 5-9 Comparison of spectra for Suite B bins of ground motion histories, 5% damping

0 0.5 1 1.5 2 2.5 3 3.5 40

0.2

0.4

0.6

0.8

1

Period (sec)

Psu

edo−

acce

lera

tion

(g) Bin 1

Bin 2Bin 3

0 0.5 1 1.5 2 2.5 3 3.5 40

0.1

0.2

0.3

0.4

Period (sec)

Psu

edo−

acce

lera

tion

(g) Bin 2

Bin 3Bin 4Bin 5

249

a. AASHTO pseudo-acceleration spectrum b. AASHTO displacement spectrum

c. Mean spectra vs. AASHTO spectral shape

Figure 5-10 AASHTO elastic response spectra with comparison to mean pseudo-accel-eration spectra for Suite B ground motions, 5% damping

0 1 2 3 40

0.2

0.4

0.6

0.8

1

1.2

Period (sec)

Psu

ed

o−

acc

ele

ratio

n (

g)


varies as 1/T

0 1 2 3 40

5

10

15

20

25

Period (sec)

Dis

pla

cem

en

t (in

)


0 0.5 1 1.5 2 2.5 3 3.5 40

0.2

0.4

0.6

0.8

Period (sec)

Psu

edo−

acce

lera

tion

(g) AASHTO

BIN motionsvaries as 1/T

250

5.4 An Evaluation of the AASHTO Uniform Load Method

5.4.1 Introduction

The Guide Specifications contain simplified procedures for estimating the design

displacement and force demand imposed on bridges designed to be seismically isolated.

The Uniform Load Method is the linearized single degree-of-freedom procedure for

estimating design response.

The Guide Specifications’ Uniform Load Method is essentially an application of the

“secant stiffness method”. The method establishes that equivalent linear properties for the

subject isolated bridge may be defined by the secant stiffness of the system responding at

its peak displacement. The method determines a design value for peak system

displacement from the ordanates of two smoothed design spectrum (adjusted for

“equivalent damping”) at the systems “secant period” (defined by its mass and “secant

stiffness”). Because the method requires an initial estimate of system properties to initiate,

the method is iterative. The method is assumed to converge to an adequate design

“estimate” of peak average time-history response of an isolated bridge system subjected to

suite of seismic events representative (on average) of the design spectrum.

5.4.2 AASHTO Uniform Load Method

The Uniform Load Method is a unidirectional method prescribed for estimating response

(i.e., forces and deformations) along each of the two orthogonal axes of an isolated bridge

system (generally chosen as the longitudinal and transverse axes). As stipulated in the

Guide Spec, total response is evaluated along each axis of the bridge by either of two

methods: the Single Mode Spectral Method or the Multi-Mode Spectral Method. The

251

Single Mode Spectral Method is focused on here.

A typical structural model of an isolated bridge bent responding uni-directionally in its

transverse isolated mode is presented in Figure 5-11 below. The force-deformation

hysteretic response of typical isolation bearings is generally broad and relatively stable.

The AASHTO Guide Specification allows isolation bearings to be characterized as bilinear

hysteretic for purposes of the design procedures as shown in Figure 5-12 (a). The overall

force-deformation response of the this typical isolated bridge system defined at the level

of the bridge deck defined in the parameters identified in the AASHTO Guide Spec,

neglecting dynamic substructure damping or mass contributions, is presented in Figure 5-

12 (b).

The code prescribes linearized characteristics for the isolated bridge system based upon

secant stiffness properties defined at its maximum displacement. “Effective damping"

properties are postulated by equating the energy dissipated per cycle in the isolated bridge

system to that of an "equivalent" linear visco-elastic damping component (oscillating

Figure 5-11 Structural idealization of an isolated bridge bent

K sub

M

di

d

keff

keff= effective isolator stiffnessKsub= linear substructure stiffnessM = lumped deck massdi = isolator displacementdsub = substructure displacementd = total displacement = di + dsub

(1

xcluding substructure contributions) (

g ratio, where rational first estimates may

), with B established from the code val

of the bridge from Equation (21).

sign spectrum closely on average and

nce). Evaluating the AASHTO Unifo

f records provides a rational basis

ing a reliable estimate of mean respo

ative values imply that the proced

fective system properties varies negligibl

quivalent static force estimates, where

o be deterministic quantities, preclud

ameter α, which represents the isolat

characterized by its idealized hyster

ange from 0 to 1. This captures isolat

Appendix E of this report. Figure

aluation of the AASHTO Uniform Lo

ons. Results plotted are the mean relat

are presented over the entire range

oad Method tended to increasingly ov

bearing properties become more flexi

rigid initial stiffness (i.e., Tsub = 0

the secant method fairly accurately

response on average for these suites

ors tended to increase with increas

re E-2, E-3, and E-4 (a)). As before,

orter characteristic periods (i.e., Tiso

he characteristic period of the isolat

s, the Method A procedure tended

ge, particularly for systems with sho

sults for the LA10in50 ground motio

omputed by Equation 22. Results

res (i.e., Τsub = 0.05 seconds) utiliz

tire range of parameters studied. For

more “nonlinearity” in the system,

nsitivity” of the error to variations

in general as the characteristic isolat

iti l l ti i d l th T

od B evaluation results for the LA10in

for the mean relative error computed

al range of parameters as the Method

(

increased with increasing characteri

o). On the other hand, βeff was redu

ncreasing Tsub), particularly for stron

f β i E ti 18 h i

Teff > 3 seconds and beff > 0.3 occur

e Method B evaluation for this suite

ratios greater than 30% were neglected

nsistent comparison to the Uniform Lo

errors for the Method B procedure.

ariation in results for the Method A

the LA10in50 database ground motio

ge systems determined by the Unifo

enerally accurate or conservative (ov

K2’

K1’K2’

K1

K2>

od” procedure to specific ground mot

re) tended to produce less conservat

study for all suites of ground moti

od” is more accurate on average w

the design setting, despite the mere f

quake events of equal likelihood poss

r” motions would be critical for isolat

for “mean” demand. These systems m

g y j p

). The earthquake demand utilized in

um (see Figure 5-6 (a)), which represe

of exceedence of 10 percent in 50 ye

w t ea t qua e ag tude a d o s

ille, 1997]. Consequently, it follows t

ectrum representing random directiv

ds represented by the mean spectrum

rcent) [AASHTO, 1999]. However, th

of elastic systems, and therefore may

f the bi-directional response of isola

d. 45o displacement path

dy

udinal and transverse axes of the bri

modeling. Five percent of the total m

gree-of-freedom. A damping ratio o

e damping contributions characterized

meters

s was performed utilizing the generali

model developed in these studies (

applied separately. Mean values of th

rectional input on system response.

ws,

motion pairs directly given the aver

motions. This can be seen clearly in

p

and isolation properties. Generally, m

asing system first-slope stiffness (

otions had more effect on systems w

stance from the earthquake fault. Figu

ratio generally decreased

esults were consistent for all ranges

xy x y

Dx Dy,( )

Values for both the Cxy coefficient

systems with the most rigid first-sl

le summarizes these statistical results

each ground motion bin. Ranges un

tudy for the ranges of ground mot

and isolated bridge system parame

of this increase due to bi-directional in

rcent respectively were realized over

This is also consistent with the results

was most pronounced for more ri

t s study d cate t at t s acto

lacement response due to bi-directio

system subjected to a suite of N grou

Dxy/mean(Dx,Dy) computed in this stu

rage value of the Dxy/mean(Dx,Dy) ra

parameters and ground motion reco

e of the D /mean(D D ) ratio ran

s of the design basis ea thquake

without its risks. For a stipulation

imply inelastic substructure behavior

ble motions However since no furt

dures provides sufficient overstrength

er, since the Guide Spec provides that

in lieu of the equivalent linear Unifo

med for certain isolated bridge syste

cte at o was ut ed.

gn Procedures

des general design criteria in Section

vides that the design of components

gy p

nd Q = M in the above design equatio

f flexural strength by dividing by an

redundancy, and ductility. Equation

of spectrum compatible ground motio

erstrength within the design proced

f the overall method. This effect will

d h t i d b li it t

ords. Isolated bridge system parame

he requirement for bounding analy

lastic. Substructure mass and damp

a d da p g co t but o assu ed to

etailed description of this model. Ove

eck is prescribed by a trilinear hystere

isolation and substructure hyster

ch earthquake record (n of N total) we

50” suite. For the case of rigid-pla

period represented by the substruct

al elastic period of the isolated bri

as generally more effected by change

Tiso) than by changes in isolator stren

ploying the most flexible and strong

t t d i th t d f t t

y demand generally increased as isolat

ond-slope stiffness, the system first-sl

co s de ed o t e 0 50 g ou

io D/Do (see Section 5.6.4 above). Th

re yielding had little effect on the to

most substructure hardening (i.e., Tsub

ere computed in this study.

ratio Cb/Cbo (see Section 5.6.4). Th

and isolation system strength decrea

hardening ratio (αsub) tended to incre

pe periods (Tsub <= 0.5 seconds) as s

indicates that the variations in syst

acement contribution from the isolat

contribution in the yielding substruct

average were computed for the majo

b ratio tended to be greatest for syste

s the first-slope period increased. Fo

ratio tended to increase with decreas

th a hardening ratio αsub 0.05. T

ith strength and second-slope flexibi

ively. Figure G-17 shows a time-hist

r the two systems in subplot (a) and

anent offset at the end of the respo

d hysteretic response results to the LA

f the lower shear output of the system

bstructure yielding. As seen in Figure

acement for the system with yield

for the system with a linear ela

ove st e gt t e des g p ov s o

mplied here. Further, the definition

so be at issue here. This may imply lo

in concrete piers, or similar limit s

espect, p ov s o s s ou d qua t y

demands and account for them explic

e incorporated through strength des

velop prior to brittle failure and throu

the response of isolated bridge syste

ustrate the effect of variations in syst

racteristics on peak response measu

Cyiso, Tiso, Tsub, respectively, see Sect

each of the three fundamental syst

em characterization was utilized in th

e AASHTO prescription could be m

o s we e pe o ed ut g t e p a

tion 5.2 and Figure 2-4(b)). Isolat

retic model mimicking the prescrib

Substructure mass was assumed to

x H of this report.

ts for the peak displacement (D (in))

field. For near-fault, fault-normal t

ed slightly with decreasing strength

ss (see Figure H-1 for Bin 1 motions

h l t ill t t th t k b h

0, and 1000 percent for Bin 1, 2, 3, 4,

udy for a variation in initial elastic per

show that peak total displacement w

nd-slope flexibility (i.e., varying Tiso)

ed to motions of larger magnitude

duced significantly with increasing T

ons (i.e., Cyiso = 0.03 subjected to Bi

0 400 percent maximum were realized

hese plots again illustrate that p

g initial system period (i.e., Tsub). A

esponse was more sensitive to variati

l h d f k

5 illustrate in more detail the effect

or systems having fixed isolator seco

t peak base shear was generally redu

ms subjected to far-field and/or sma

ted to ground motions at greater dista

thquake magnitudes (see Figure H-1

for isolated bridge systems subjected

at a given distance range from the act

idge systems with longer initial ela

base shear with decreasing earthqu

stems with a fixed second-slope per

her from the active fault (i.e., Bin 3 an

ely flat along the x-axis of the p

d decrease markedly along the positiv

yiso).

h H-26 present mean contour results

lustrate that substructure displacem

elastic period (i.e., increasing Tsub) fo

cond-slope period (i.e., longer Tiso,

e flexibility for these ground motion

ucture displacement for an increase

ith fixed isolator second-slope flexibil

ent reduced with increasing strength.

ame less effective in reducing isola

d over the long period range. The spec

placement was dependant on the spec

gures H-28 through H-31). The larg

etail the effect of isolator second-sl

ment response for systems having fi

ator displacements were generally l

n substructure displacements.

y p g

e that peak isolator displacement did

increasing initial elastic period. Furth

displacement was generally reduced

f earthquake distance on peak isola

k isolator displacement was reduced

und motions at greater distance from

f magnitudes). This reduction in isola

riod increased.

earthquake magnitude on substruct

mean peak substructure displacement w

ti f ll th k it

ectively) occur concurrently with p

amic assumptions imposed in this stu

utions), it is clear from Equation 44 th

cross the substructure.

strate in more detail the effect of stren

os (Diso/D and Dsub/D) for systems w

ement between the isolation system

bridge may be significantly effected

itial elastic periods. For these syste

reduce the percentage of displacem

more effect on near-fault than far-fi

e H-52 through H-55 for Bin 2 throug

n D /D and D /D shown here (

strated dependencies in the response

of system properties and ground mot

g p g

nce. In this study, a reduction in stren

crease in average total displacement

shortest initial elastic periods conside

m

g

e to variations in isolator second-sl

rger isolator second-slope flexibility

A maximum reduction in average b

hed earlier). For displacement dema

the sensitivity of base shear to variati

ex. However, as increasing strength w

n these studies it is apparent that stron

ess at large displacement amplitude

Displ.

larger δForcen strength

t a e ast c pe od dec eased).

acement increased in general as isola

an due to variations in strength over

tor second-slope flexibility increased

g ca t y as t e t a e ast c pe od

substructure flexibility increased). T

s shown earlier, where since isola

engthened or reduced with increas

os

articipating mass or damping, peak to

olator and substructure displacement

er implies that for these assumptions

ture and base shear response, it was a

structure displacement distribution w

ponse (characteristic of lower magnit

motions) and more sensitive to seco

t uctu e d sp ace e t at o educed)

lt distance. The change in response

ed was dependant on the specific syst

ion.

y g p

ution illustrated a similar, but not

econd-slope stiffness, most sensitive

s, distribution of displacements betw

be effected by varying the second-sl

f a significant effect on total and isola

ective as response amplitude decrea

a linearized method for estimating

ated bridge systems (see Section 5.4

structure approach facilitates simplif

ethod and/or other alternative procedu

ibutions to response.

m substructure components were mode

the substructure degree-of-freedom

f the substructure damping element, c

ue to time constraints, half of the grou

very other record for each bin select

computed, with the difference being t

bridge system with and without a gi

ain computed as D/D0, Diso/Diso 0, D

s for the peak displacement ratio (D/D

) which compare the peak displacem

substructure damping of ζsub = .05 to

ping contributions in a similar mann

and that for Bin 1 motions is due to

r displacement demands (as compared

of strength variations imposed at l

t of substructure damping of ζsub = 0

e was essentially negligible in all ca

y 3% maximum. In the response to ne

amping of ζsub = 0.05 on peak total

0.05) was 40% and 70% on avera

ubjected to Bin 5 ground motions.

ar Response

n 5 motions. For these higher frequen

asing system strength tended to incre

r initial elastic periods and decrease

form sensitivity to either strength

5.7.

so show that the Dsub/Dsub 0 ratios w

ess effected (compare data from Table

ally essentially rigid (i.e., initial ela

mping of ζsub = 0.05 suppressed p

se a maximum of approximately 25%

s are on the order of 5-30%, as sho

997]). Results from Chapter 2 sugges

he D/Do ratios were smallest for stron

smallest for weaker systems (with Tsu

eld) motions were uniformly effected

eriods considered (i e T b = 0 05 t

Do ratio was generally less than unity

sed (i.e., γ = .05 to .10). For Bin 1 (ne

over the entire range of initial ela

. This indicates for these cases that

ve and/or below unity for the mean D

es indicate the most significant variat

.974 1.08 1.11.99 1.11 1.2

1.11 .999 1.18 1.16 1.2

1.04.99 .999 .984 1.06 1.

s in isolator second-slope flexibility (

ratio to increase as Tiso increased wh

e systems were more sensitive to hig

ystems.

ues in Figure I-11 to I-15.

/or below unity for the mean Diso/Dis

es indicate the most significant effect

.873 .919 .839 1.04.947 .86

1.29 1.28 1.42 1.25 1.4

1.38 1.26 1.5 1.51 1.5

o ratios were effected predominantly

early insensitive to system strength (

became more dependant on strength

ure mass ratio increased from γ = .05

stribution is most significant in

ubstructures.

te to these values. These effects

study imply that ignoring higher m

ed with the distribution of substruct

contributions may cause signific

es regarding the relative significance

of these studies are outlined below

nd base shear response is generally m

than peak total and isolator displacem

d generally less effected for systems w

m Table 5-22 through 5-25) For syste

or other ground motion bins, the effec

and isolator displacements became m

flexible (i.e., Tsub >= 0.5 seconds), w

5 to 60 percent for substructure m

these contributions must reasonably

y is needed to establish whether curr

ation or other approximate methods

s to response, without resorting to m

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6 Conclusions and Recommendations

6.1 Introduction

This report presented the results of a series of experimental and analytical studies

conducted for the California Department of Transportation under the Protective Systems

Research Program at the University of California, Berkeley. This report covers research

developed in order to:

1. Establish an understanding of global and localized response characteristics of simpleseismically protected bridges subjected to various types of seismic input including far-field, near-fault, and soft-soil motions.

2. Evaluate the effect of bi-directional and three-dimensional loading on isolated bridgesystems.

3. Validate the efficacy of seismic protective systems for bridges of simple, but realisticconfigurations.

4. Assess the applicability of requirements in the AASHTO Guide Specifications andrecommend improvements where needed based on the results of this research.

Each of these issues was examined earlier in Chapter 2 through 5 of this report. Reference

each of these chapters directly for more complete information on specific results. In this

concluding chapter, these findings are summarized and used to assess current design

procedures, and develop recommendations for improvement where necessary.

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With respect to these needs in the current state-of-practice, the results of this study provide

two major conclusions.

• Foremost, earthquake simulation tests of system performance developed under this researchhave provided crucial proof-of-concept for the application of protective systems technology tobridges. These tests and ancilliary analyses have illustrated the efficacy of seismic isolation forsimple single- and double-span overcrossings and more complex bridge segments subjected tomulti-dimensional inputs of far-field, near-fault, and soft-soil ground motions. Key resultsfrom these earthquake simulation studies and analyses are summarized below to illustrate thispoint (see Sections 6.2 and 6.3).

• Secondly, analytical evaluations performed in this research have contributed significantly tothe understanding of the behavior of isolated bridge systems and insight into appropriatedesign approaches. These analytical studies have also established several areas in the currentprovisions of the Guide Specifications in need of further development.

Key analytical results are presented below in detail, along with recommendations for

improvement of code procedures where appropriate (see Sections 6.3).

6.2 Earthquake Simulation Studies

6.2.1 Proof-of-concept

The earthquake simulation studies developed in this project (see Chapter 3) verify the

efficacy of seismic isolation of simple bridge overcrossings. These tests illustrate the

robustness of bridge isolation as a means to mitigate damage during a seismic event.

Numerous input signals representing design basis far-field, near-fault, and soft-soil ground

motion inputs were imposed on the bridge specimens developed for these studies.

Insignificant variation in performance was realized after multiple and repeated tests.

Configurations with significant mass and stiffness eccentricity also showed effective

isolated response to numerous and varied ground motion input. These tests also illustrated

the effect of substructure mass, substructure yielding, vertical motions, and secondary

geometric nonlinearity on the response of these systems. Analytical trends identified in

Chapter 2 showing the effect of basic system variations on response were essentially

375

validated by these experimental simulations. Key results from these simulations are

summarized below.

6.2.2 Sensitivity to characteristics of ground shaking

The effect of variations in ground motion input on the effectiveness of isolated response

was illustrated in these simulation studies. Far-field, near-fault, and soft-soil events were

considered in these studies as well as the effect of multi-dimensional coupling in these

motions (i.e., the correlation of uni-directional, bi-directional, and vertical motion

components).

6.2.2.1 Ground motion type: far-field, near-fault, and soft-soil

Three earthquake table motions (i.e., LA13_14, NF01_02, and LS17c_18c) were selected

for studies of the effect of variations in ground motion type on the response of simple

isolated bridge systems. Four substructure stiffnesses represented by the various bridge

specimen configurations allowed this evaluation to be carried out over a broad range of

spectral frequencies.

Bridge specimen test results to these ground motion inputs indicated that response was

strongly influenced by ground motion characteristics, including frequency content and

directional orientation. The LA13_14 motion had significant higher frequency content and

response was strongly oriented along a 45 degree axis to the x-y direction. This resulted in

specimen response with significant coupling in the x- and y-directions and underlying

cycling of higher frequency. The LS17c_18c motion, on the other hand, was a soft-soil

motion with long period content near the characteristic period of the test bearings and

produced relatively harmonic response in the specimen. The near-fault NF01_02 motion

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had impulsive, strongly coupled, fault-normal and fault-parallel components and produced

strongly coupled behavior in the specimen.

In general, the response of the bridge specimen could be classified based on the corellation

of its initial elastic period with the ranges of spectral peaks for LA13_14, NF01_02, and

LS17c_18c records. To illustrate this point, versions of the x-direction component of each

of three earthquake table motions were processed at two length scale factors (i.e., lr). In

this way a structure would respond as if it were in a different intensity and spectral

frequency range. This extended the range of structure dynamic characteristics considered

in these studies.

These analytical results indicated that for isolated systems with very rigid substructures

(or shorter initial elastic periods) peak total displacement is much larger than that of a

similar elastic system (i.e., non-isolated bridge with the same mass and initial stiffness

characteristics). For systems with greater substructure (or initial) flexibility, total peak

displacement was roughly equal to that of a similar elastic system. This behavior was

characteristic of systems in these “short period” and “long period” (or equal displacement)

ranges of spectral response (see also preliminary analytical results of Chapter 2). The

period ranges defining these “short” and “long” period ranges was a characteristic of the

particular ground motion input, as seen in the response to the scaled versions of these table

motions (see Chapter 4). For example, the “long period” (equal displacement) range for

the LA13 table motion began around Tsub = 0.55 seconds for a scale factor of lr = 1/2,

while for a length scale factor of lr = 1/4 the “long period” range began around Tsub = 0.35.

To illustrate the variation in response characteristics for systems in these different spectral

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ranges, consider for example an isolated bridge system with a fixed fundamental (or initial

elastic) period of Tsub = 0.4 seconds. This isolated system may be in the “short period”

range (i.e., exceeding elastic displacements) for the LA13 motion at lr = 1/4. For the same

motion at lr = 1/2, however, this system would be in the “long period” range (i.e., having

nearly equal displacement as an elastic system). Response characteristics to different

ground motion input can be evaluated similarly.

6.2.2.2 Bi-directional motions

The effect of bi-directional input on specimen response was evaluated by computing the

two ratios and from test results. Here Uxy is the

peak displacement response from the bi-directional test while Ux and Uy are the peak

displacement from the x- and y-direction uni-directional tests, respectively. Results of

these ratios for the various specimen configurations indicate that the effect of bi-

directional input is most pronounced for systems with the most rigid initial stiffness (see

Figure 4-12). Values of Uxy/mean(Ux,Uy) and Uxy/srss(Ux,Uy) covered the range of 1.4-1.8

and 0.9-1.2 on average, respectively, for systems with the most rigid substructures (see

Chapter 4). For comparison, values of Uxy/mean(Ux,Uy) = 1.414 and Uxy/srss(Ux,Uy) = 1.0

would indicate that peak Ux and Uy displacements were occurring simultaneously in the

bi-directional response, resulting in peak Uxy displacement equal to their vector sum.

Further in-depth analytical studies on these effects were carried out in Chapter 5.

Uxy mean Ux Uy,( )⁄ Uxy srss Ux Uy,( )⁄

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6.2.2.3 Vertical motions

Experimental tests considered the effect of vertical motions on the response of a bridge

specimen utilizing a sliding isolation system. Test data indicated that vertical input tended

to effect isolator and total displacement response more than substructure displacement and

isolator shear force response. Peak isolator and total displacement response were

increased by approximately 9 and 5 percent on average by the addition of the vertical

input component, while peak substructure displacement and isolator shear force were

effected by less than 3 and 1 percent on average, respectively. The characteristic strength

of FP bearings fluctuate more with axial load variation than their stiffness, since Qd = µN

and kd = N/R and µ is generally an order of magnitude larger than 1/R. Consequently, since

deck and isolator displacements were shown to be more sensitive to system strength than

substructure displacements and isolator forces (see Chapter 2 and 5) the effect on them is

more pronounced.

6.2.3 Effect of system configuration

The effect of variations in basic system properties on isolated response was also studied in

these simulation studies. System variations included substructure mass, strength and

flexibility; isolator strength and flexibility; and global mass and stiffness eccentricity.

6.2.3.1 Substructure flexibility

The various configurations of the bridge deck specimen incorporated substructures with a

range of flexibility. Subsequent analyses on isolated bridge systems with mechanical

properties similar to the bridge deck specimen were also carried out for comparison with

these test results.

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In these experimental and analytical results peak deck displacements for isolated systems

with very rigid substructures were much larger than those of similar non-isolated systems

that are strong enough to remain elastic. Nearly all the displacement in the isolated system

occurs in the isolator component. On the other hand, for isolated systems with relatively

large substructure flexibility total displacement was roughly similar to that of a similar

non-isolated system. For these more flexible systems, total displacement was shared in

varying proportions between the isolation bearings and the substructure component

(dependant upon the characteristics of the ground motion input). Peak deck displacement

was seen to monotonically increase with increasing substructure flexibility. Peak isolator

displacement also increased, but less dramatically, or for certain inputs was roughly

constant for systems varying over the same range of substructure flexibility.

Peak deck acceleration response for these isolated systems was seen to be relatively

constant over a range of substructure flexibility. By comparison, for a similar non-isolated

elastic systems peak acceleration response would undulate dramatically as system

flexibility was varied (dependant again on characteristics of the ground motion input).

The advantage of the isolated system therefore is not significant force reduction, but

damage control and enhanced performance. Performance is enhanced as the isolation

system limits force response and makes it less sensitive to input characteristics. Damage is

controlled as the isolated system endures significant displacement demands (taken in a

large portion through the isolation bearings) without yielding or damage to the

substructure. Conventional bridges, on the other hand, are designed to endure significant

substructure damage in a design basis earthquake event.

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6.2.3.2 Isolator strength and flexibility

The configurations of the bridge deck specimen incorporated isolation system with

different mechanical properties for comparison.

For the isolated bridge specimen with an essentially rigid substructure, results indicated

that displacement response was particularly sensitive to strength and systematically

increased as strength decreased. Base shear response, on the other hand, was particularly

sensitive to isolator second-slope stiffness and consistently increased as bearings second-

slope stiffness increased. In addition, the stiffening behavior of HDR bearings caused by

scragging effects at peak (virgin) response cycles resulted in further increases in shear

response above those experienced by other bearings at similar displacement amplitudes.

Notably this increased force spike in the HDR bearings near peak amplitude did not

appear to markedly change displacement response.

For the isolated bridge specimen with more flexible substructures (i.e., Tsub = 0.25, 0.75

and 1.0 seconds), displacement response was also sensitive to strength and systematically

increased as strength decreased. Strength was seen to have more effect on displacements,

however, as the substructure became more rigid. This was more apparent at larger motion

amplitudes. Force response, on the other hand, was less sensitive to variations in

characteristic isolator strength. Variations in isolator second-slope flexibility were not

evaluated on these specimen configurations.

6.2.3.3 Substructure mass

Varying substructure mass proportions of γ = msub/Mdeck equal to approximately 0, 5, and

10 percent were considered in the simulation studies. Results indicated that the effect of

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variations in substructure mass over this range was most prominent on substructure

displacement (and consequently base shear), less on isolator displacement, and least on

total deck displacement response. However, the difference in average response (to the

three table test motions) as substructure mass was varied from 0 to 10 percent, rarely

exceeded 10 percent. The tendency was for average response to increase as substructure

mass increased. Base shear response was therefore highest for the system with most

substructure mass (resulting directly from the higher substructure displacement demand),

as expected. On the other hand, additional substructure mass had a relatively insignificant

effect on isolator shear force response. This behavior is intuitive, since increased isolator

displacement results in only minor increase in force output due to the low second-slope

stiffness of these devices.

6.2.3.4 Mass eccentricity

Superstructure mass eccentricity of 0, 5, and 10 percent of span length was considered in

these studies for a simple single-span bridge overcrossing. Limited test results indicated

that torsional contributions to peak isolator displacement at the ends of the deck span were

more prominent for impulsive and soft-soil type motions than for far-field events. Further,

isolator displacements at the ends of the deck span did not, in general, increase linearly

with mass eccentricity.

Previous evaluations suggest FP bearings provide a re-centering mechanism to an isolated

system [Bozzo, et al 1989]. The friction force and lateral stiffness of an FP bearing is

proportional to the applied normal force (see Chapter 3, Equation 3). This implies the

center of resistance of an isolated system utilizing FP bearings will inherently align with

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the system’s center of mass, counteracting torsional response. Analysis and testing in

these previous studies confirmed this behavior [Bozzo et al, 1989].

The tests performed in this report using impulsive motions, however, seem to refute this

generalization. This may be due to the sensitivity of “rigid” structures to strength (see

Sections 2.3.5.2 and 4.3). The lighter end of the specimen (having a lower resisting

friction force) may tend to displace more than the heavier (i.e., stronger) side producing

the torsional behavior seen in these studies. Comparable results to test this hypothesis for

full intensity motions utilizing LR bearings (whose strength is not sensitive to axial

applied force) were not possible in these studies. In this respect, further analysis and

testing is still needed to establish reliable trends in the response of systems with eccentric

mass.

6.2.3.5 Stiffness eccentricity

The effect of substructure stiffness eccentricity on isolated response was also evaluated in

these studies. Significant stiffness eccentricity was considered with the braced and

unbraced ends of the specimen differing in total lateral stiffness by an order of magnitude

(i.e., nearly ten times). Configurations utilizing FP bearings with symmetric slider

arrangements and unsymmetric arrangements arranged to counteract torsional response

(with weaker sliders on stiffer braced end piers and stronger sliders on flexible end piers)

were also considered.

These limited results indicated that peak rotation demand in the eccentric configurations

was not a function of peak displacement amplitude nor was it effected significantly by bi-

directional input. Prior to yielding in the isolation bearings, stiffness eccentricity induced

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significant torsional response. However, when bearings operate in their isolated mode on

either end of the specimen overall system stiffness eccentricity is minimized. The rotation

response in these eccentric configurations was therefore due to the tendency of the system

to rotate prior to bearing yield. Post-yield (i.e., in the isolated mode) stiffness eccentricity

is nearly eliminated and rotation demands become insensitive to incremental and peak

displacement response. Isolated behavior in essence suppresses the underlying torsional

behavior, unlike in a standard bridge where substructure stiffness eccentricity would

impose torsional demands increasing with response amplitude.

Further, bi-directional input in these tests imposes an additional orthogonal component of

motion on the longitudinal axis of the specimen. In this direction the stiffness arrangement

was symmetric. Thus, the additional longitudinal component did not engage lateral-

torsional coupling in the system, and results indicated that this additional component did

not tend to increase rotation demands significantly. Coupling in the yield response of the

isolation bearings imposed by the additional longitudinal component did not appear to

effect peak rotation significantly either in these tests. This again indicates that torsional

response in these systems, with substructure stiffness eccentricity, appears to be more

effected by system behavior prior to bearing yield.

Tests results indicated that peak total displacement response was somewhat larger for the

same bridge specimen with added substructure stiffness eccentricity (i.e., one end braced

rigidly). Peak isolator displacement response was not as systematically effected

(sometimes increasing and sometimes decreasing), indicating that torsional rotation was

not consistently in phase with peak response in these test sequences. These limited results

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also indicated an unsymmetric configuration utilizing lower and higher strength FP sliders

installed in bearings located above braced and unbraced supports, respectively, was

effective in reducing torsional response due to underlying substructure stiffness

eccentricity.

6.2.3.6 Substructure yielding

The effect of yielding substructure components on the response of a two-span isolated

bridge specimen illustrated a pattern of force redistribution from central yielding piers to

outer (non-yielding) end piers. Despite this yielding behavior, total deck displacement

history remained stable during the isolated response. In this system, isolation bearings

absorbed the majority of total displacement demand up to yielding in the substructure.

Beyond this point, central piers yielded plastically until peak displacement was attained.

As such, only minor ductility demands were imposed on the substructure. For a

conventional (non-isolated) bridge system, nearly the entire displacement demand would

be imposed on the elements of the substructure, inducing significantly larger ductility

demands on these components.

6.2.4 Characteristic distribution of force and displacement demands

Typical patterns of global and local force and displacement distribution were noted in

these studies. Shear demands tend to concentrate toward stiffer end (abutment) piers.

Center pier-isolator assemblies see similar total displacement demand as end pier-isolator

assemblies enforced by displacement compatibility in relation to the in-plane stiffness of

the bridge deck as it displaces laterally. Since the secant stiffness of the central pier-

isolator assemblies at peak displacement is lower typically than at end (abutment) pier-

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isolator assemblies, higher shear demands concentrate at abutment locations.

Displacement distribution between isolator and pier components are made in relation to

the relative flexibilities of these elements at their respective locations. Since each location

sees nearly equal total displacement demand, a smaller proportion of the total

displacement at center piers is taken in the isolation bearings with a larger proportion

taken in the more flexible center piers. On the other hand, nearly all the displacement

demand is taken by isolation bearings at stiffer end (abutment) piers.

6.2.5 Local kinematic effects

In two-span bridge specimen tests, end (abutment) piers were braced while center piers

could rotate on clevis pins at their base. Consequently, during testing flexible center piers

underwent significant lateral tip displacement and overall rigid body rotation. This

rotational behavior caused second-order axial shortening in the vertical height of these

central piers. Axial shortening in stiffer braced end (abutment) piers was insignificant.

Axial shortening in center piers resulted in redistribution of gravity load to end piers

transmitted by the stiff continuous girder.

Since the slip-force and lateral stiffness of FP bearings are proportional to the supported

axial load (see Chapter 3, Equation 14), the vertical force redistribution in these test

sequences just described caused a significant pinching of the force-deformation behavior

of the FP bearings atop central piers. Globally, total axial force (due to gravity) remained

constant, thus rendering total bearing slip-force and total system strength unchanged.

Locally, however, the redistribution of axial force to end piers, which increased bearing

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strength and stiffness locally, increased shear demand on associated connections and

substructure components at these locations.

This particular phenomenon appeared to be more a characteristic of the unique test setup

in this case then that representative of an actual bridge. However, axial load redistributions

will occur in real bridge structural systems to some extent. These tests highlight the

sensitivity of sliding and FP bearings to fluctuations in axial load, effects that may impose

local force amplifications. The shear stiffness of elastomeric bearings is also effected by

vertical stress variations, and their stability at peak displacement is determined on the

basis of vertical load demands.

The second order kinematic effects causing axial load redistribution in these test

sequences are not typically incorporated in standard design analysis methods nor included

in current design procedures [AASHTO, 1999]. These results highlight the need for

further research in this area. The effect of axial load fluctuations on sliding systems due to

vertical ground motion input was discussed previously, above.

6.3 Implications for Design Practice

Current design practice of isolated bridge systems is embodied in the AASHTO Guide

Specifications for Seismic Isolation Design [AASHTO, 1999]. Results of the preliminary

evaluations presented in Chapter 2, simulation tests presented in Chapter 3 and 4, and the

analytical studies presented in Chapter 5 have indicated several areas where application of

current design procedures is adequate and pinpointed other areas in need of further

consideration. Recommendations for specific improvements are offered below. In some

cases, alternate design methodologies or approaches are suggested. These methodologies

387

may effectively mitigate some of the difficulties encountered in the application of the

current Guide Specifications. Further research in some areas is also needed, and several

recommendations for future research are outlined as well.

6.3.1 Basic AASHTO Design Equation

6.3.1.1 Reliability

The guide specifications are based on approximate analytical methods that linearize the

response of inelastic systems. Many general studies of such methods have been under

taken. Most of these linearization methods are based on either harmonic or random

response [Chopra and Goel, 1999]. It has been shown in numerous studies that methods

based on harmonic response considerably overestimate the period shift of the substitute

system, whereas methods derived considering random response give much more realistic

estimates of effective period [Iwan and Gates, 1979b]. Two methods based upon harmonic

response have been widely adapted to the design several types of inelastic structures. The

"substitute structure method" [Shibata and Sozen, 1976] has been popularized by some for

displacement-based design [Gulkan and Sozen, 1974; Shibata and Sozen, 1976; Moehle,

1992; Kowalsky et al., 1995; Wallace, 1995]. The "secant stiffness method" [Jennings,

1968] has been adapted to formulate the "nonlinear static procedure" in the ATC-40

[Applied Technology Council, 1996] and FEMA-274 reports [FEMA, 1997]. For the

design of seismically isolated bridges, the AASHTO Guide Specifications [AASHTO,

1999] utilizes the Uniform Load Method which is an adaption of the "secant stiffness

method".

388

Since the Uniform Load Method is based on an assumption of ideal harmonic response,

there is a concern that it may overestimate the shift of the system’s effective period [Iwan

and Gates, 1979b]. In other words, the Uniform Load Method may tend to overestimate

the system’s “effective” flexibility, thereby overestimating displacement response.

Concern has also been raised about the Uniform Load Method’s linear visco-elastic

“equivalent” damping assumption, as this might not be appropriate for response to

strongly impulsive motions (such as those characteristic of near-fault shocks). Dissipation

provided by viscous damping is rendered less effective under impulse excitations.

In these research studies, it was found that the displacements of seismically isolated bridge

systems computed by the Uniform Load Method were generally accurate or conservative

(over-predicted) on average over a broad range of simple isolated bridge structures for a

suite of spectrum compatible ground motions. This suggests that period shift was

generally predicted accurately or over-predicted by the procedure. This is consistent with

previous findings [Iwan and Gates, 1979b]. On the other hand, direct application of the

underlying “secant stiffness method” procedure in these studies using specific ground

motion records tended to still produce conservative, but less conservative, estimates of

peak response on average. This suggests that the “secant method” was more accurate on

average when applied to individual ground motion spectrum directly then when applied to

a smoothed representation of the mean spectrum of a suite of motions (as in the Uniform

Load Method procedure). The level of conservatism is not uniform and some ranges of

structural characteristics produce very conservative results that could have a significant

impact on project costs.

389

It was also found that the difference in errors produced by application of the “secant

stiffness method” to far-field, near-fault, and pure pulse type ground motions were

inconclusive in determining whether the harmonic damping assumptions utilized in the

linearization procedure are less conservative for more impulsive ground motion types.

Other structural factors and the relation of ground motion and structural characteristics

were found to be more important as factors in determining the accuracy of the method

than the ground motion characteristics alone.

The basis of the Uniform Load Method on an assumption of linearized behavior suggests it

would necessarily produce more accurate prediction for more “linear” systems. This is

achieved as either kd and ku become equal (i.e., the isolator becomes more “linear”), the

substructure becomes more flexible (i.e., Ksub decreases), or as system strength decreases

(i.e., Qd is reduced). System “linearity” may be evaluated through the parameter αsys (see

Chapter 5, Equation 27), where the system becomes more “linear” as αsys approaches 1.0

(see Chapter 5, Figure 5-13). Strength reduction increases system “linearity” by both

decreasing system “effective damping” and causing system second-slope and secant

stiffness to converge (see Chapter 5, Equation (18) and (26)). Trends in the error between

the peak response computed by the Uniform Load Method and the “exact” response

computed by nonlinear time-history analysis in these studies appear to validate this

prediction. The Uniform Load Method was most accurate for systems with lower strength,

having more flexible substructures, or utilizing more “linear” isolation devices (i.e., kd

approaching ku).

390

The results of this study also suggest the inter-relationship between system strength and

first-slope stiffness effects the accuracy in the Uniform Load Method procedure. For more

initially rigid systems (i.e., highest first-slope stiffness), errors in Uniform Load Method

were reduced as system strength reduced consistent with the above discussion (i.e., the

system became more “linear”). However, for more initially flexible systems (i.e., lower

first-slope stiffness) errors were reduced as strength increased. This trend likely resulted

from the tendency for stronger systems to behave more predominantly in their linear first-

slope mode.

The Guide Specifications present limits on the application of the Uniform Load Method

presumably to prohibit usage of its linearized procedure for more nonlinear systems.

Explicit nonlinear analysis is required both for systems with long equivalent periods (i.e.,

Teff > 3 seconds) and for systems which are highly “damped” (i.e., beff >= 0.3, unless B =

1.7 is used). In these studies, these measures did not appear to mitigate error uniformly,

however, with both accurate and inaccurate results alike being discarded by the

application of these limits.

Notably, when the Uniform Load Method was most inaccurate in this study, it tended to

overpredict response on average. However, these results also suggest significant under-

estimation may occur for ground motions with spectra above the mean design spectrum.

Assuming a normal distribution in the statistical scatter, mean-1σ errors as large as 50-75

percent were computed for the suite of twenty motions considered in these studies. These

twenty motions were selected to match the design spectrum closely on average and

possessed the same return interval as the target design spectrum. It is important to note

391

that all the earthquake events in this suite are equally as likely to occur within the

representative return interval. Demands from the “larger” motions in the suite would

therefore be critical for isolation systems designed by the Uniform Load Method for

“mean” spectral demand. These systems may be provided with limited overstrength and/or

surplus displacement capacity to resist these larger demands. Similarly, substructure

components designed to remain “essentially elastic” for “mean” design response by the

Uniform Load Method may not possess adequate overstrength or ductility capacity to

resist the level of demand imposed by these larger events. In the end, design based on

“average” spectral intensity masks the risk associated with these larger demands. It may

prove appropriate for the Guide Specifications procedures to rationally account for the

probability of these less numerous, but equally likely, events.

The results of this study showed that the Uniform Load Method had non-uniform

accuracy, suggesting that some systems will be designed more conservatively and some

less conservatively by these procedures. Further, the Guide Specifications limit on

equivalent period and damping (i.e., Teff 3 seconds and beff 0.3) did not resolve this

issue in a consistent fashion. Therefore, while this study has found that the method

appears to provide a level of conservatism on average, alternate displacement-based

design methods may be warranted to achieve more uniform reliability. Measures which

allow rational consideration of the likelihood of risk associated with larger events of

equivalent return period should also be incorporated. A discussion of alternate procedures

which incorporate these features is presented later.

≤ ≤

392

6.3.1.2 Bi-directional effects

The AASHTO Guide Specifications define earthquake demand through a smoothed design

spectrum representing a uniform earthquake hazard with a probability of exceedence of

10% in 50 years defined regionally across the nation (see Chapter 5, Figure 5-6 (a)). It is

understood that real ground motions consist of shaking along two horizontal axes and a

vertical axis at a given site. The AASHTO design spectrum represents the uniform hazard

spectrum for expected ground motion records. As such, it represents random or “average”

directivity. With regard to estimating the peak displacement of isolated bridges subjected

to bi-directional inputs, the use of a design spectrum representing random directivity to

estimate response demands poses two distinct difficulties.

First of all, it has been illustrated that the difference between the mean spectrum

representing fault normal or the “larger” orthogonal component of bi-directional ground

motion history pairs and the mean spectrum of all ground motion history pairs

(representing “average” directivity) increases with earthquake magnitude and for sites

located more closely to the active fault [Somerville, 1997]. Consequently, it follows that

design displacement estimates based upon a spectrum representing random directivity

would underestimate mean displacement demands represented by the mean spectrum of

the “larger” (fault normal) orthogonal component of the earthquake record. This

underestimation would be most severe for larger magnitude earthquakes and near-fault

sites.

Secondly, it is not apparent that the peak response of a seismically isolated bridge system

subjected to bi-directional input can be estimated with reasonable accuracy from the

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response of the same system subjected to each of the bi-directional components applied

uni-directionally. The reason for this is twofold. Primarily, phasing within each of the

ground motion components may dictate bi-directional response with peaks occurring more

or less simultaneously producing a maximum vectored displacement much larger than

either of the uni-directional maxima. Next, coupling which occurs in the bi-directional

yield surface of typical seismic isolation systems causes a reduction in resisting force

orthogonal to the direction of initial displacement (see Chapter 5, Figure 5-14). This

coupling effect would presumably cause increases in displacement demand for systems

subjected to bi-directional motions, since displacement along one axes will reduce the

isolation system’s resistance to motion in the orthogonal direction.

Current code procedures provide a method of estimating peak bi-directional force demand

in isolated bridges by combining uni-directional maxima in a combination of 100 to 30

percent (or conversely 30 to 100 percent) [AASHTO, 1999]. However, these procedures

were established from the evaluation of elastic systems, and therefore may not adequately

capture the nonlinear complexities of the bi-directional response of isolated bridge

systems.

An evaluation of the effects of directivity and bi-directional input on the response of

simple seismically isolated bridge systems was undertaken in these studies. It is evident

from the results presented that the effect of bi-directional input on seismically isolated

bridge systems is significant. It was found for all cases considered that on average peak

displacement response due to bi-directional input was considerably larger than peak

response due to uni-directional input applied separately. This disparity was most

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significant for more rigid structures, employing stronger isolation systems, subjected to

larger magnitude earthquakes, and located nearer to the active fault. For these cases, softer

site-specific soils tended to further increase the bi-directional effect.

These results are consistent with the previous factors which influence the effects of bi-

directional coupling. Firstly, coherent phasing within each of the ground motion

components has been shown to be most significant for near-fault motions (which contain a

more coherent impulsive content than in the far-field) and for soft soils. The results of this

study are consistent with this trend, where it is seen that the bi-directional effect is more

pronounced for these ground motion conditions. Second, coupling in the force-

displacement yield surface of seismic isolation bearings would tend to be more

pronounced for more rigid systems, where interaction with the yield surface would occur

more readily at smaller displacements. In addition, the greatest reduction in resisting force

orthogonal to the direction of initial displacement would occur for the strongest isolation

systems. This is also consistent with the results of these studies, where the bi-directional

effect was most pronounced for systems which were stronger and more initially rigid.

As described previously, the AASHTO Uniform Load Method provides a procedure for

estimating the mean peak displacement response of seismically isolated bridge systems

subjected to a suite of uni-directional spectrum compatible motions representing random

or “average” directivity. If it is assumed that this method provides a reliable estimate of

the mean peak response of an isolated bridge system subjected to a sufficiently large

sampling of spectrum compatible motion pairs, then the results of this study indicate that

the method will significantly underestimate on average the peak displacement response of

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the system when subjected to the same ground motion pairs applied bi-directionally. In

addition, since the difference between the mean spectrum of “larger” (fault-normal)

ground motions and the mean spectrum of all ground motion pairs (representing “average”

or random directivity) increases with closer proximity to the active fault (due to directivity

in the ground motion), the AASHTO Guide Specifications procedures have been shown in

these studies to be increasingly less conservative in accounting for bi-directional effects

with closer fault proximities or soft-soils. The Guide Specifications allow a combination

of 100 percent plus 30 percent of orthogonal maxima to be utilized to account for the

effects of bi-directional input [AASHTO, 1999]. If this procedure is applied a maximum

vectored displacement of only approximately 1.04 would be realized, assuming design

response in each orthogonal direction to be equal. The results of this study indicate that

this factor is inadequate to capture average increases in displacement response due to bi-

directional coupling.

On average, in these studies increases in peak displacement response due to bi-directional

input of ground motion pairs of approximately 25-75 percent above the average response

due to the same pair of ground motions applied separately (i.e., uni-directionally) were

realized for the ranges of ground motion distance, magnitude, site specific soil type, and

isolated bridge system parameters considered (see Chapter 5, Table 5-18). It is therefore

recommended that a revision to code procedures be applied to account for the effects of bi-

directional interactions. For simple bridge overcrossings, substructure stiffness is typically

dictated by essentially rigid abutment conditions. Further, design basis earthquake demand

is nearly equivalent to the Bin 2 motions considered in these studies (see Chapter 5).

Given these general conditions for design, it is recommended that the effects of bi-

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directional motions for simple bridge overcrossings be computed by weighting design

displacements as follows

(50)

where dbi is the peak bi-directional displacement in any vectored direction, Cxy is a bi-

directional weighting factor, and dx and dy are the uni-directional design displacements

computed by Guide Specifications procedures in each orthogonal direction of the bridge,

respectively. It is recommended that a factor of Cxy= 1.5 be used for the weakest and most

initially flexible isolation systems (where bi-directional interactions would be least) and

Cxy= 1.7 be used for the strongest and more initially rigid isolation systems (where bi-

directional interactions would be greatest). For near-fault and softer site specific soil

conditions these factors should be increased by an additional 5%.

6.3.1.3 Equivalent SDF vs. MDF behavior: Influence of Substructure Mass and Damping

The linearized substitute structure procedure employed in the AASHTO Uniform Load

Method is a conceptually appealing approach which facilitates analysis. However, it has

limitations with regard to the incorporation of substructure mass and damping

contributions. These contributions make the determination of “equivalent” single-mode

“effective stiffness” and “effective damping” properties difficult. The Guide

Specifications allows a Multi-Mode Spectral Method to be utilized to incorporate the

effects of non-isolated higher modes which may potentially capture substructure mass

effects. Alternative damping procedures for accounting for substructure contributions

have also been suggested by others [Sheng et al., 1994]. Ignoring these contributions is

dbi Cxydx dy+

2-----------------

=

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also a practical simplification. However, possible loss of accuracy in the methodology

may result. To this end an evaluation of the effects of substructure mass and damping

contributions on the response of simple seismically isolated bridge systems was

undertaken in these studies. These results suggest that substructure mass and damping

contributions play different roles in effecting total, isolator, and substructure response.

The magnitude of these influences may determine when these contributions can be

neglected in the design process without significant loss of accuracy, or on the other hand,

when ignoring these contributions may cause significant underestimation of peak

response.

The results of these studies illustrated that for isolated bridge systems with essentially

rigid substructures the effect of adding substructure damping of 5 percent of critical (i.e.,

ζsub = 0.05, see Chapter 5, Equation 49) on peak total and isolator displacement response

was negligible in all cases (suppressing average response by less than approximately 3

percent maximum). For these systems (i.e., simple bridge overcrossings with essentially

“rigid” abutments), these substructure damping contributions could be reasonably

neglected. For response to near-fault motions (i.e., Bin 1), the effect of 5 percent

substructure damping on peak total and isolator displacement response was minor even for

systems with the most flexible substructures, with a peak reduction in average response of

approximately 10 percent maximum. For far-field and lower magnitude earthquakes (i.e.,

Bin 2 through 5), the effect of 5 percent substructure damping on peak total and isolator

displacement response became increasingly more significant with increasing substructure

flexibility. For these motions the addition of 5 percent substructure damping produced

maximum reductions in average peak total and isolator displacement of 40 and 70 percent,

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respectively, for the most flexible substructures considered (i.e., Tsub = 2 seconds). In all

cases, the addition of substructure damping caused reductions in average peak total and

isolator displacement response. Ignoring these contributions would be considered a

conservative simplification (as they result in reductions in computed average peak

displacement). For systems with more flexible substructures which may benefit

significantly from these reductions, this simplification may be perhaps considered overly

conservative.

These results also indicated that for systems with stiffer substructures the addition of 5

percent substructure damping had more effect on average peak substructure displacement

and base shear response than on peak total and isolator displacement. On the other hand,

for systems with more flexible substructures the opposite was true. For systems with

essentially rigid substructures (i.e., Tsub = 0.05 seconds), 5 percent substructure damping

reduced average peak substructure displacement and base shear response by

approximately 25 percent maximum in these studies (compared to 3 percent for total peak

and isolator displacement response). For systems with the most flexible substructures

considered (i.e., Tsub = 2 seconds), 5 percent substructure damping reduced average peak

substructure displacement and base shear response by approximately 30 percent

maximum (compared to 40 and 70 percent for total and isolator displacement,

respectively). While these reductions in substructure displacement and base shear

response may be considered significant, ignoring these reductions would again be

considered a conservative simplification in the design process. Particularly, since it is the

performance goal in an isolated bridge to limit substructure damage, considering higher

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force demands for these components (by ignoring substructure damping contributions)

will contribute to this result.

The effect of typical substructure mass contributions of 5 and 10 percent of the total

system mass (i.e., γ = .05 to .10, see Chapter 5, Equation 48) were examined in these

studies as well. Results indicated for systems with essentially rigid substructures that 5

and 10 percent substructure mass contributions reduced average peak total and isolator

displacement response approximately 8 and 16 percent maximum, respectively. For

response to near-field, motions (i.e., Bin 1), 5 and 10 percent substructure mass

contributions had only a minor effect on peak total and isolator displacement response for

all ranges of substructure flexibility considered (with response varying by approximately

10 percent maximum). For these cases, ignoring substructure mass contributions may be

considered a reasonable simplification (as they result in reductions or otherwise minor

variations in total and isolator response). For other ground motion conditions (i.e., Bins 2

through 5), the effect of substructure mass contributions up to 10 percent on peak total and

isolator displacements became more significant as substructure flexibility increased, with

average response increased by approximately 15 to 60 percent over the range of isolation

system parameters considered. For these cases, ignoring substructure mass distribution

may result in significant underestimation in computed response, particularly for systems

with more flexible substructures.

The results of this study showed that substructure mass contributions always increased

average peak response of the substructure. For systems which were initially essentially

rigid (i.e., Tsub = 0.05 seconds), substructure mass contributions of 5 and 10 percent

400

increased peak substructure response approximately 50 and 90 percent maximum,

respectively. For systems with the most flexible substructures (i.e., Tsub = 2 seconds), 5

and 10 percent substructure mass contributions increased average peak substructure

response approximately 10 and 25 percent maximum, respectively. Systems with

intermediate substructure flexibilities, showed increases in response between these values.

These effects are significant, and ignoring these substructure mass contributions would

likely cause considerable underestimation of average peak substructure response (i.e.,

displacement and base shear).

These results are useful in establishing the significance of substructure mass and damping

contributions to the global response of simple isolated bridge systems. Conditions in

which these contributions may be reasonably ignored without significant loss of accuracy

have been identified in the discussions above. Ignoring these contributions in these cases

is a practical simplification. For cases where more accurate consideration of substructure

damping is warranted, alternative equivalent damping procedures which seek to account

for these contributions are available. Many of these methods have been examined by

others [Sheng et al., 1994]. Strain-energy proportional methods for determining these

effects have also been presented [Kawashima et al., 1994]. For cases where more accurate

consideration of substructure mass contributions are warranted, the Guide Specifications

allows the Multi-Mode Spectral Method to be utilized to account for these higher mode

effects. Further study is needed to establish whether this procedure is adequate in

compensating for these contributions. More accurate modeling discretization and explicit

time-history analysis may be required in some cases to account for these higher modes.

401

6.3.1.4 Parametric Study of Nonlinear Isolated Bridge Response

Due to the underlying difficulties with current linearized Guide Specifications procedures,

and the need to understand the trend in nonlinear displacement and force response of

isolated bridges for purposes of design, a systematic evaluation of the nonlinear response

of seismically isolated bridge systems was undertaken in these studies. This parametric

evaluation was developed to establish the specific effects of pertinent design variables on

the response of isolated bridge systems. These fundamental variations include mechanical

properties of the isolation system and substructural components as well as variations in

features of the earthquake ground motion input.

For purposes of this study, a bilinear simplification was utilized to characterize isolation

bearing hysteresis and total system uni-directional force-deformation response (see

Chapter 2, Figure 2-2 (a) & (b)). Three independent parameters (eg., the first- & second-

slope stiffness, K1 & K2 respectively, and the system yield strength, Fy) define this

bilinear system. An evaluation of the influence of these hysteretic parameters on system

response was used to understand the explicit nonlinear behavior of simple isolated bridge

systems.

Total Displacement. - These studies illustrated that total displacement response was very

sensitive to variations in the initial elastic period of the isolated bridge system. Total

displacement response increased as much as tenfold on average in these studies for an

increase in initial first-slope elastic period from 0.05 to 2 seconds for systems with fixed

strength and isolator second-slope flexibility.

402

Total displacement response was also found to be more sensitive to variations in system

strength than isolator second-slope flexibility over the range of properties considered (i.e.,

strength coefficeients from 3 to 12 percent and bearing characteristic isolation periods

from 2 to 6 seconds). In this respect, increased strength resulted in decreased total

displacement for systems with shorter initial elastic periods and increased total

displacement for systems with longer initial elastic periods. The spectral region at which

strength had negligible effect on average total displacement became centered at

progressively longer periods for ground motions of increasing magnitude and decreasing

fault distance.

Similar dependency on strength had been reported previously in the nonlinear response of

single-degree-of-freedom bilinear oscillators subjected to earthquake and shock inputs

[Newmark, et al. 1975]. For these inputs imposed on systems with initial elastic periods

less than the predominant pulse period, displacement increases dramatically as strength is

reduced. For systems with initial elastic periods near the predominant pulse period,

however, displacement decreases with reduced strength. This behavior was illustrated

earlier in Chapter 5, Figure 5-16. The total displacement response seen in these studies

mimics this reported behavior.

Base Shear. - These studies illustrate that base shear response was less sensitive than total

displacement to variations in initial elastic period. Maximum decreases in mean base shear

of 50 percent occurred in these studies for an increase in initial elastic period of 0.05 to 2

seconds over the range of isolation system parameters and ground motion inputs

considered.

403

Generally, base shear decreased as strength was reduced and/or isolator second-slope

flexibility was increased. However, base shear was more sensitive to strength for low

amplitude response (characteristic of Bin 3, 4, and 5 motions) and more sensitive to

isolator second-slope flexibility for large amplitude response (characteristic of Bin 1

motions) in these studies. For system response in the intermediate range of amplitudes

(characteristic of Bin 2 motions), weaker systems with higher isolator second-slope

stiffness had base shear response more sensitive to variations in isolator second-slope

flexibility while stronger and/or systems with larger isolator second-slope flexibility had

base shear response more sensitive to strength.

This trend in base shear behavior may be explained by considering fundamental system

response at varying displacement amplitude ranges. For lower magnitude and/or far-field

motions in this study (characteristic of Bin 3, 4, and 5 motions), total displacement

response was relatively small. As shown earlier for small amplitude displacement

response (see Chapter 5, Figure 5-17 (a)), variations in system second-slope stiffness have

a relatively minor effect on overall force output (and on displacement response as well, as

discussed previously above). Comparatively, an increase in system strength, although

perhaps reducing overall displacement, will also produce a significant increase in overall

force output. On the other hand, for near-fault, fault-normal motions in this study (i.e., Bin

1 motions), total displacement demands were relatively large. As shown earlier for this

type of large amplitude response (see Chapter 5, Figure 5-17 (b)), variation in system

strength, although perhaps effecting overall displacements dramatically, does not have a

relatively large impact on force output compared to the effect of system second-slope

flexibility imposed over this larger displacement demand (understanding that

404

displacements are not significantly effected by variation in second-slope flexibility, as

established earlier). For displacement demands intermediate to these cases, as for Bin 2

motions, the sensitivity of base shear to variations in strength and isolator flexibility is

more complex. However, as increasing strength was seen to reduce displacement response

generally in these studies, it is apparent that stronger systems subjected to these motions

will have displacement demands toward the smaller amplitude range with base shear

response showing sensitivity as shown in Chapter 5, Figure 5-17 (a). Weaker systems

subjected to these motions, on the other hand, will have displacement demands toward the

larger amplitude range with base shear response showing sensitivity as illustrated in

Chapter 5, Figure 5-17 (b). However, for systems with the largest second-slope flexibility

(i.e., larger Tiso) response is nearly elastic perfectly-plastic. These systems will have force

output necessarily more sensitive to changes in strength as shown earlier in Chapter 5,

Figure 5-17 (c).

Isolator Displacements. - These studies illustrate that isolator displacements were more

sensitive in general to system strength than isolator second-slope flexibility, similar to

total displacement response. Further, increased isolator strength decreased isolator

displacements consistently on average in these studies. For example for a strength increase

from 3 to 12 percent, maximum reductions in isolator displacement of 80 percent in

isolator displacement occurred for systems with the shortest initial elastic periods in these

studies.

These results also illustrated that isolator displacement increased in general as isolator

second-slope flexibility increased, but less so than due to variations in strength over the

405

range considered. Sensitivity to changes in isolator second-slope flexibility was greatest

for weaker, stiffer isolation bearings and for systems with longer initial elastic periods.

On the other hand, unlike total displacement response isolator displacements did not

increase or decrease in general with increasing initial elastic period. For far-field motions,

average isolator displacement response in this study increased by approximately 100

percent but also decreased by as much or more as initial elastic period increased from 0.05

to 2 seconds for specific isolation system parameters over the range of properties

considered. For near-fault motions, isolator displacement in this study reduced nearly 40

percent maximum with increasing initial elastic period for systems with the weakest and

most flexible isolation systems considered, but in general remained relatively constant.

Substructure Displacement. - For these studies, isolation systems were considered rigid

plastic such that initial system flexibility was attributed entirely to the substructure

component. These studies illustrated that substructure displacement was dependant upon

its stiffness. Substructure displacement response increased significantly as the initial

elastic period of the isolated bridge system increased (i.e., as substructure flexibility

increased), similar to total displacement response. Since isolator displacement response

was seen to neither consistently increase or reduce with increasing substructure flexibility,

it is implied that increase in total displacement is largely the result of increasing

substructure displacement as the substructure is made more flexible.

Furthermore, these studies illustrated that substructure displacement showed sensitivity to

strength for low amplitude response and to second-slope isolator flexibility for large

amplitude response similar to base shear. Similarly for intermediate amplitude response,

406

substructure displacement was most sensitive to strength when employing the strongest

and most flexible isolation systems and most sensitive to isolator second-slope flexibility

when employing the weakest and stiffest isolation systems.

Finally, these studies illustrated that displacement of flexible substructures was the most

sensitive to variations in isolation system properties. In general, substructure displacement

response was reduced by reducing strength and/or increasing isolator second-slope

flexibility.

Isolator and Substructure Displacement Ratios. - These studies illustrated that in general

the ratio Diso/D decreased (and inversely Dsub/D increased) on average as substructure

flexibility increased (i.e., increasing initial elastic period) for all isolation system

parameters and ground motion inputs considered. In addition, isolator and substructure

displacement distribution illustrated the same sensitivity to strength and isolator second-

slope flexibility as base shear and substructure displacement response (for a system with a

fixed substructure stiffness). With this in mind, these studies illustrated that reduced

strength increased the isolator’s contribution to total displacement and reduced it for the

substructure (i.e., increased Diso/D and decreased Dsub/D) for a system with fixed initial

stiffness. Increasing isolator second-slope flexibility caused a similar redistribution.

Earthquake Magnitude and Distance. - These studies illustrated for all cases of system

parameters considered that larger earthquake magnitude and/or nearer fault distance

increased total displacement, isolator displacement, substructure displacement, and base

shear response on average over the suites of ground motion records utilized in these

407

studies. In addition, the isolator displacement ratio was increased (and the substructure

displacement ratio reduced) for larger earthquake magnitude and/or nearer fault distance.

Further, as system response amplitude increased at larger earthquake magnitude and/or

nearer fault distance the sensitivity of base shear, substructure displacement, and

substructure to isolator displacement distribution changed from strength to isolator

second-slope flexibility. Increasing response amplitude did not effect appreciably the

sensitivity of total and isolator displacement to strength, however.

Summary. - Performance objectives in an isolated bridge system typically include control

of force and displacement demands. The results of these studies are useful in establishing

the causal links in ground motion and system characteristics effecting these performance

measures.

Total and substructure displacement response were shown in these studies to be

particularly sensitive to initial elastic system stiffness. On the other hand, base shear

response was shown to be relatively insensitive to this variation. In addition, isolator and

total displacement response was shown to be more sensitive to strength and relatively

insensitive to variations in isolator second-slope flexibility. Further, base shear and

substructure displacement response illustrated a dependency on strength and isolator

second-slope flexibility in relation to the amplitude of response (i.e., more sensitive to

strength for low amplitudes and more sensitive to second-slope flexibility for the large

amplitude response). Isolator and substructure displacement distribution illustrated a

similar, but not as dramatic, dependency on strength and isolator second-slope stiffness,

most sensitive for flexible substructures.

408

Taking these dependencies into consideration, response may be tailored to meet specific

performance objectives and design situations. Practical limits of isolation system

mechanical properties and substructure component proportioning must be considered in

this respect, as well as the potential seismic environment for the proposed application.

Reduction in total and/or substructure displacements for an isolated bridge system may be

most effected by an increase in initial elastic system (or substructure) stiffness. These

studies indicate that this will have a minor effect on total base shear and likely a minor

effect on isolator displacement (although in the far-field, a likely reduction in this as well).

Isolator displacements are best reduced by adding strength, although adding initial elastic

stiffness may also provide a significant reduction in these displacements (although may

also perhaps produce an undesirable increase).

Base shear (and substructure displacements) can best be reduced for lower amplitude

response (i.e., in the far-field) by reducing strength, anticipating an associated increase in

total and isolator displacements. Reducing base shear (and substructure displacements) by

adding isolator second-slope flexibility is also an option, although most effective for

larger amplitude response (i.e., in the near-field). Further, since total and isolator

displacements are relatively insensitive to this variation, this latter method is promising

for reducing base shear without producing an over-compensation in displacement

response.

For systems with relatively flexible substructures, distribution of displacements between

the isolator and substructure component can be effected by varying the second-slope

flexibility of the isolation system (without fear of a significant effect on total and isolator

409

displacements). This method becomes less effective as response amplitude decreases.

Varying isolator second-slope flexibility should be executed with discretion, as this

measure may either unduly increase base shear response or produce large residual

displacements which may require post-earthquake recentering.

In addition, for systems with relatively flexible substructures displacement distribution

between the isolator and substructure can also be effected by strength. This measure

becomes less effective as response amplitude increases. Also, re-distributing

displacements in this manner will also effect isolator and total displacements significantly,

requiring this measure to be balanced with other performance objectives.

The understanding of these trends in the nonlinear behavior of isolated bridges is key to

removing the dependence on linearized assumptions which cloud current design

procedures and provide limited guidance to designers. Nonlinear spectra of the types

produced in this study produced from a large statistical sample could also be adapted

directly for use as a design methodology. This would provide a measure of rationality in

the procedure and would allow uncertainty to be statistically quantified in the design

method. Falling short of this, utilization of the understanding of explicit nonlinear

behavior developed here in conjunction with current linearized design procedures will

assist designers in proportioning simple isolated bridge systems to meet desired

performance.

Finally, current Guide Specifications procedures require consideration of variations in

isolation device properties in the design procedure. This is instituted by requiring

component properties to be multiplied by λ factors which have been empirically

410

formulated from test results to bound mechanical properties. Results of the parametric

studies undertaken in this research lend further clarity to these measures. These results

indicate that variations in certain properties may produce negligible variation in system

response and thus their consideration is irrelevant in the bounding analyses. On the other

hand, particular variations significantly effect system response (e.g., strength). In these

cases, any additional rational measures aimed at considering lower and upper bounds in

these properties are certainly well justified.

6.3.1.5 Substructure yielding

The formulation of the AASHTO Guide Specifications suggests that some measure of

substructure yielding may occur. This is implied by stipulating the use of force reduction

factors which eliminate overstrength in the demand-capacity equations for substructure

design for the case of an isolated bridge (see Chapter 5, Section 5.1.1). The Guide

Specifications commentary states that these lower R-Factors (in the range of 1.5 to 2.5) are

formulated such that their “ductility based portion is near unity and the remainder

accounts for material overstrength and structural redundancy” and design by these

provisions “ensure, on average, essentially elastic substructure behavior in the design-

bases earthquake”. The Guide Specifications commentary further states that response

calculated by the prescribed procedures “represent an average value, which may be

exceeded given the inherent variability in the characteristics of the design basis

earthquake” [AASHTO, 1999]. However, since no further guidance is given in the Guide

Specifications on how to assess or provide for this yielding behavior, the formulation

implies that yielding of substructural components will not adversely effect the

411

performance of the isolation system nor inflict ductility demands on substructure

components beyond their inherent limits.

On this basis, studies were performed to evaluate the efficacy of this new code

formulation in its ability to meet these stated performance objectives (i.e., ensure elastic

behavior on average). The results of these studies indicate that significant yielding of

substructure components may occur for isolated bridge systems designed optimally to

these current AASHTO design provisions. Most notably, substructure ductility demands

greater than 1.0 on average were computed in these studies in all cases (and in many cases

much greater than 1.0). This appears to refute the claim that the provisions in the Guide

Specifications “ensures essentially elastic substructure response on average” [AASHTO,

1999]. Either way, it appears columns in isolated bridge systems designed by Guide

Specifications provisions will require a measure of ductile capacity (since ductility

demands are at least implied by the commentary for response above the mean). However,

since a significant portion of the overall displacement demand occurs in the bearings for

an isolated bridge system, ductility demands in the columns of isolated bridges will still be

substantially smaller than in a comparable non-isolated system.

Further research is needed to establish whether the AASHTO detailing provisions provide

sufficient nominal ductility capacity in their component design specifications to meet the

demands implied by this study. The definition of “essentially elastic substructure

response” may also be at issue here. This definition may imply local strain demands to the

level of surface spalling in concrete piers, or similar limit state definitions, which would

412

be produced at a level of global ductility demand larger than unity. Further research is

needed to address these ancillary issues.

The results of these studies also indicate that the presence of substructure yielding in

systems designed optimally to the AASHTO Guide Specifications had negligible effect on

overall total and isolator displacement demands, but instead caused a redistribution of

displacement between the isolation system and the yielding substructure component. This

effect may limit displacement demands on isolation devices and in turn drive displacement

ductility demands into the substructure. Further, when substructure yielding is explicitly

allowed, shears are reduced - a desirable effect which will limit forces on foundations,

connections, and brittle components. However, the effects of isolator overstrength due to

variations in isolator material properties, contamination, aging, scragging, or path effects;

impact on displacement restraint systems; and system vs. component overstrength

relations may result in increases in force output and produce larger substructure ductility

demands on systems designed without their consideration. These additional factors must

be considered in any rational design procedure.

Finally, if the intentions of the AASHTO provisions are to ensure elastic substructure

response on average, then it is implied that ductile response will be allowed for 50 percent

of “design basis” ground motions. In this respect, provisions should quantify the

magnitude and risk associated with these implied demands and account for them explicitly

in component design procedures. Given a determination of these substructure ductility

demands, strength design procedures may then be utilized to ensure ductile mechanisms

413

develop prior to brittle failure and that proper detailing requirements are supplied to

provide sufficient plastic rotation, local strain, and/or inelastic buckling capacity.

Barring a more detailed procedure, a general procedure for estimating substructure

ductility demands in simple bridge overcrossings may be postulated as follows:

1. Since total displacements were shown in these studies to be generally conserved,substructure and isolator displacement demands may be estimated for simple isolatedbridge overcrossings given that

(50)at peak response. This simplification may be considered appropriate for bridgeovercrossings since the effect of substructure mass and damping contributions onisolator and substructure displacements was shown in these studies to be essentiallynegligible for all but the most flexible substructures.

2. Isolator and substructure component nonlinear force-deformation relations may thenbe established though test results, mechanical relationships, or as for column bentsthrough standard moment-curvature analysis procedures.

3. Demands for isolator and substructure elements may then be established to satisfy theabove relation assuming force compatibility for these components at computeddisplacements.

4. As a final step, ductility demands for substructure components may be evaluatedthrough the established force-deformation relations from the previous step.

6.3.2 Alternative Methods

Displacement based design methods are currently utilized as standard practice. These

procedures focus on establishing displacement demands (and associated forces) for the

purpose of system design. Examples of these are incorporated in the ATC-40 [Applied

Technology Council, 1996] and FEMA-274/356 [FEMA, 1997] building codes. The

AASHTO Guide Specifications for Seismic Isolation Design (or Guide Spec) [AASHTO,

1999] also incorporates displacement based procedures for the design of isolated bridges.

Diso Dsub+ D≈

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Displacement based methodologies can be loosely characterized into equivalent linear and

approximate nonlinear procedures. More recently performance-based design (PBD)

methodologies have been undergoing development as well. PBD methods go beyond

earlier approximate procedures and seek to rationally quantify the likelihood of specific

limit or loss state outcomes given the conditional probability of system response in a given

seismic environment. A summary of equivalent linear, approximate nonlinear, and PBD

procedures are presented below.

6.3.2.1 Equivalent Linear Procedures

Equivalent linear analytical methods for estimating response of inelastic systems include

methods based on either harmonic or random response [Chopra and Goel, 1999].

Currently, two methods based upon harmonic response have been adapted to the design of

inelastic structures. The "substitute structure method" [Shibata and Sozen, 1976] has been

popularized by some for displacement-based design [Gulkan and Sozen, 1974; Shibata

and Sozen, 1976; Moehle, 1992; Kowalsky et al., 1995; Wallace, 1995]. The "secant

stiffness method" [Jennings, 1968] has been adapted to formulate the "nonlinear static

procedure" in the ATC-40 [Applied Technology Council, 1996] and FEMA-274/356

reports [FEMA, 1997]. For the design of seismically isolated bridges, the AASHTO Guide

Specifications [AASHTO, 1999] has adopted the Uniform Load Method which is

essentially an adaption of the "secant stiffness method".

It has been shown that methods based on harmonic response considerably overestimate

the period shift of the substitute system, whereas methods derived considering random

response give much more realistic estimates of effective period [Iwan and Gates, 1979b].

415

Further, it was shown in these studies that the Uniform Load Method utilized in the

AASHTO Guide Specifications does not produce results of uniform reliability. Results of

these studies indicate that this method may produce slightly over-conservative designs for

certain system characteristics and slightly under-conservative designs in other cases (see

Chapter 5, Section 5.4).

6.3.2.2 Approximate Nonlinear Methods

Approximate nonlinear methods have been pioneered by numerous authors. Earlier

techniques focused on establishing system strength capacity necessary to limit global

ductility demands [Newmark, 1975]. These techniques focused on establishing strength

reduction factors which permit estimation of inelastic strength demands from elastic

strength demands. A comprehensive review of various investigations of strength reduction

factors carried out over the last thirty years has been performed by others [Bertrero, 1994].

Recently nonlinear displacement based procedures have been formulated as well. These

methods focus on establishing inelastic displacement ratios defined as the ratio of the

maximum lateral inelastic displacement demand of a structure with given strength to the

maximum lateral elastic displacement demand for a system with similar first-slope

stiffness properties [Miranda, 2000]. Researchers have reviewed several of these new

procedures and compared advantages and disadvantages of each method [Miranda, 2002].

Approximate nonlinear displacement based procedures have also been formulated for

standard practice. Examples of these are incorporated in the ATC-40 [Applied Technology

Council, 1996] and FEMA-274/356 [FEMA, 1997] building codes.

416

While both forced-based and displacement-based approximate nonlinear methods are

useful, and explicitly recognize the nonlinear behavior of real structural systems subjected

to earthquake overloads, their formulation has been predominately developed for building

structures which inherently have a limited upper bound on ductility. Isolated systems

characteristically respond with displacement ductilities significantly greater than those

imposed on non-isolated building systems. Therefore, currently available approximate

nonlinear procedures have limited use for design of isolated bridges without significant

reformulation to consider larger nonlinear displacement amplitudes.

6.3.2.3 Performance-Based Design

Performance based design (PBD) methods seek to enable accurate probabilistic

quantitative evaluation of structural performance. These methods integrate consideration

of random variations in input and system characteristics and evaluate their effect on

nonlinear response to discern statistical bounds on likely outcomes.

Probabilistic models. - The development of PBD methods has focused initially on the

formulation of probabilistic models of system performance. Recently a comprehensive

Bayesian methodology for developing probabilistic capacity and demand models for

structural components and systems has been formulated [Der Kiureghian, Mosalam, et al,

2002]. These probabilistic models are similar to deterministic capacity and demand

procedures commonly used in practice, but apply additional correction terms to account

for inherent systematic and random errors. These models provide means to gain insight

into the underlying behavioral phenomena and to select ground motion parameters that are

most relevant to the seismic demands. The models take into account information gained

417

from engineering principles, laboratory test or field data, and engineering experience and

judgment [Der Kiureghian, Mosalam, et al, 2002].

PEER methodology. - The Pacific Earthquake Engineering Research Center (PEER) is

engaged in developing a complete probability-based framework for performance-based

design. This formulation is both rational and adaptable to a variety of applications. The

methodology comprises several models including a seismic demand model. The PEER

Center is not alone in this endeavor. Both FEMA (the Federal Emergency Management

Agency) and ASCE (the American Society of Civil Engineers) have developed the

FEMA/ASCE 356 prestandard. This document addresses performance in terms of facility

operability, occupiability, life safety, and resistance to collapse, under four discrete levels

of seismic excitation. PEER's methodology seeks to address economic performance as

well as operability and safety, and to express performance in probabilistic terms such as

distributions on repair costs and loss-of-use on an annualized or lifetime basis.

PEER's performance based design methodology is illustrated schematically in Figure 6-1

below. This PBD formulation seeks to quantify the likelihood of specific limit or loss state

outcomes rationally conditional on system response in a given earthquake hazard

environment. The methodology incorporates hazard analysis and structural analysis, both

familiar aspects of current design practice. The PEER methodology also adds two new

features, damage and loss analysis. Damage analysis is the explicit, probabilistic

calculation of physical damage (eg., which bars have buckled, which beams have spalling,

etc.) for a given level of global demand. Loss analysis is the explicit, probabilistic

418

calculation of performance in terms of economic loss and loss of use for a given level of

damage.

PEER's performance-based methodology is currently in development. It is being evaluated

in the PEER testbed project, which seeks to demonstrate and exercise the methodology on

six real facilities: two buildings, two bridges, a campus of buildings, and a network of

highway bridges. In conjunction with this, engineering practitioners are comparing the

PEER methodology with current practice (such as the ASCE/FEMA prestandard and other

approach), to identify strengths and areas of needed development [Deierlein, 2001].

Bridge design parameter sensitivity. - Currently researchers are utilizing the development

of recent probability-based PBD framework to develop probabilistic demand models for

highway bridge overpasses [Stojadinovic, 2002]. This framework seeks to develop

demand models which relate ground motion intensity measures (IM), such as peak

Figure 6-1 PEER PBD Analysis Methodology [Deierlein, 2001]

419

spectral acceleration, to bridge demand parameters (DP), such as displacement ductility.

These models are used to assess response sensitivity to variations in bridge design

parameters, such as column height and diameter. Relations for each design parameter give

bridge designers the ability to evaluate the effect of their design choices on structural

performance [Stojadinovic, 2002].

Conclusion. - Performance based design methods are the current trend in design of

systems for earthquake performance. These methods seek to enable accurate probabilistic

quantitative evaluation of system performance and quantify decision making parameters

in terms of damage and loss. These methods integrate the consideration of variability in

ground motion input, system mechanical characteristics, and nonlinear response. These

methods are rational and adaptable and may be readily formulated for use in the design of

isolated bridge systems.

Performance-based formulations also seek to remove the dependence on approximate

equivalent linear and nonlinear procedures. There focus is to determine response based

upon accurate probabilistic distributions, and establish sensitivity for each design

parameter giving designers the ability to evaluate the effect of their design choices on

structural performance. This formulation is similar to the evaluations performed in these

studies, where earthquake hazard was defined by a suite of deaggretized motions,

nonlinear response determined explicitly through time-history analyses to these inputs,

and response presented statistically for each bridge system-input variation to illustrate

sensitivity. This PBD approach provides statistical quantification of the likelihood of

specific performance limit outcomes given system response to a given seismic

420

environment. The performance-based methodology removes the reliance on approximate

methods which cannot inherently provide the rational measure of reliability essential to

defining risk. In this respect, PBD procedures represent the next step in the development

of design procedures for earthquake resistant systems.

6.4 Future Research Need

Through the research presented in this paper, developed under the Coordinated Protective

Systems Program for the California Department of Transportation, much insight has been

gained regarding the application of these technologies to bridge systems. In particular, as

outlined in the program objectives (see Chapter 1, Section 1.3), this research has improved

knowledge in the following areas:

1. Understanding the effects of bi-directional loading on seismic isolation bearings and indeveloping improved analytical bearing models.

2. Understanding global and local response characteristics of simple seismicallyprotected bridges subjected to various types of seismic input including far-field, near-fault, and soft-soil motions.

3. The efficacy of seismic protective systems for simple and more complex (realistic)bridge configurations.

4. The efficacy of requirements in the AASHTO Guide Specifications for SeismicIsolation Design.

Results and conclusions presented in the preceding six chapters highlight this

understanding. In conjunction with this research, however, areas of future research need

have also been identified. Further development in these areas is deemed critical to assist

the application of protective systems to bridge structures to gain broad acceptance and use.

These research areas include:

1. Effect of deck flexibility and overturning on global and local isolated response.

421

2. Effect of abutment pounding on energy dissipation and isolated response.

3. Effect of bridge joints.

4. Effect of bridge skew.

5. Kinematic effects on axial load distribution.

6. Effect of random variations in isolation system characteristics.

7. Performance-based design procedures.

8. Performance hybrid systems incorporating isolators and supplemental energydissipation devices.

These areas of research are logical extensions of the studies already performed under

Phase I of the Protective Systems Research Program presented in this report.

6.5 Conclusion

The devastating consequences of a major seismic event was illustrated during the Loma-

Prieta, Kobe, Northridge and more recent earthquakes. Protective systems provide a

practical method for protecting life and property in the event of a major earthquake. The

studies performed in this research have illustrated the general ability of protective systems

to provide an effective means of earthquake resistance in simple bridge systems. The

considerable durability and robustness of these systems were illustrated through multiple

and varied simulations. Given this validation, broad acceptance of this technology in

bridge applications should be considered a priority in zones of high seismic risk. To this

end, support of new and existing protective systems applications is essential, as well as

support for continuing research efforts and the current trend of development of

performance based design procedures.

422

Bibliography

American Association of State Highway and Transportation Officials (AASHTO)."AASHTO LRFD Bridge Design Specifications (1994)", American Association of StateHighway and Transportation Officials, Washington, D.C., 1994.

American Association of State Highway and Transportation Officials (AASHTO). "GuideSpecifications for Seismic Isolation Design (1999)", American Association of StateHighway and Transportation Officials, Washington, D.C., 1999.

Applied Technology Council (1996); "Seismic evaluation and retrofit of concretebuildings", Report ATC 40, November 1996.

Bertero, V. V.; Miranda, E. (1994). “Evaluation of strength reduction factors forearthquake-resistant design”, Earthquake Spectra, 10, 2, May 1994, pages 357-379.

Bozzo, L., Zayas, V. A.; Low, S.; Mahin, S. A. (1989). “Feasibility and performancestudies on improving the earthquake resistance of new and existing buildings using thefriction pendulum system”, UCB/EERC-89/09, Berkeley: Earthquake EngineeringResearch Center, University of California, Sept. 1989.

Chopra, Anil K. (1995). "Dynamics of Structures - Theory and Applications to EarthquakeEngineering", University of California at Berkeley, Berkeley, CA, 1995.

Chopra, A. K., and Goel, R.K. (1999). "Capacity-demand-diagram methods for estimatingseismic deformation of inelastic structure: SDF system", Report: PEER 1999/02, PacificEarthquake Engineering Research Center, University of California at Berkeley, Berkeley,April 1999.

Constantinou, M.C., Mokha, A., and Reinhorn, A.M. (1990). “Experimental study andanalytical prediction of earthquake response of a sliding isolation system with a sphericalsurface”, NCEER-90-0020, National Center for Earthquake Engineering Research,Buffalo, N.Y., Oct. 11, 1990, 1 vol.

Constantinou, M.C., Reinhorn, A.M., and Nagarajaiah, S. (1998). “Nonlinear dynamicanalysis of base isolated structures - current techniques,” SEI-ASCE 13th Conference onAnalysis and Computation, paper reference T125-3.

423

Deierlein, G. G. (2001). “Peer Methodology Testbeds”, http://www.peertestbeds.net/,Pacific Earthquake Engineering Research Center, University of California, Berkeley.

Der Kiureghian, Armen; Mosalam, Khalid M.; Gardoni, Paolo (2002). “Probabilisticmodels and fragility estimates for bridge components and systems”, PEER-2002/13,Berkeley: Pacific Earthquake Engineering Research Center, University of California, June2002.

FEMA (1997). "NEHRP guidelines for the seismic rehabilitation of buildings", FEMA273; and "NEHRP commentary on the guidelines for the seismic rehabilitation ofbuildings", FEMA 274, Federal Emergency Management Agency, Washington, D.C.,October 1997.

Fenves, G. L. (1998). “Analysis and testing of seismically isolated bridges under bi-axialexcitation”, Proceedings of the 5th Caltrans Seismic Research Workshop: CaliforniaDepartment of Transportation Engineering Service Center, Sacramento, California, June1998.

Gulkan, P., and Sozen, M.A. (1974). "Inelastic responses of reinforced concrete structuresto earthquake motions", Journal of the American Concrete Inst., 71, 12, pp604-10.

Iwan, W.D., and Gates, N.C. (1979a). "Estimating earthquake response of simple hystericstructures", Journal of the Engineering Mechanics Division, ASCE, 105, EM3, pp391-405.

Iwan, W.D., and Gates, N.C. (1979b). "The effective period and damping of a class ofhysteretic structures", Earthquake Engineering and Structural Dynamics, 7, 3, pp199-212.

Jennings, P.C. (1968). "Equivalent viscous damping for yielding structures", Journal ofthe Engineering Mechanics Division, ASCE, 94, EM1, pp103-116.

Kawashima, K., Nagashima, H., Iwasaki, H. (1994). “Evaluation of Modal Damping Ratioof Bridges based on strain-energy proportional method”, Journal of StructuralEngineering, 1994.

Kawashima, K., Shoji, G. (1999). “Nonlinear hysteretic interaction between device andpier in a isolated bridge”, Proceedings of the Second World Conference on StructuralControl, John Wiley & Sons, Chichester, England; New York, Vol. 2, 1999, pp 877-884.

Kowalsky, M.J., Priestley, M.J.M, and Macrae, G.A. (1995). "Displacement-based designof RC bridge columns in seismic regions", Earthquake Engineering and StructuralDynamics, 24, 12, pp1623-43.

Krawinkler, H.; Alavi, B. (1998). “Development of improved design procedures for nearfault ground motions”, Proceedings SMIP98 Seminar on Utilization of Strong-MotionData; Oakland, California, September 15, 1998; California Division of Mines andGeology, Sacramento, California, 1998, pages 21-41.

424

Krawinkler, H.; Moncarz, P. D. (1982). “Similitude requirements for dynamic models”,Dynamic Modeling of Concrete Structures SP 73-1, American Concrete Inst., Detroit,Michigan, 1982, pages 1-22.

Miranda, Eduardo (2000). “Inelastic displacement ratios for displacement-basedearthquake resistant design”, 12th World Conference on Earthquake Engineering, NewZealand Society for Earthquake Engineering, Upper Hutt, New Zealand, 2000, Paper No.1096.

Miranda, Eduardo; Ruiz-Garcia, Jorge (2002). “Evaluation of approximate methods toestimate maximum inelastic displacement demands”, Earthquake Engineering &Structural Dynamics, 31, 3, Mar. 2002, pages 539-560.

Moehle, J.P. (1992). "Displacement-based design of R/C structures subjected toearthquakes", Earthquake Spectra, 8, 3, pp403-427.

Newmark N. (1959). "A method for Computation for Structural Dynamics", Journal ofthe Engineering Mechanics Division, 85, EM3.

Newmark, N. M. (1975). “A response spectrum approach for inelastic seismic design ofnuclear reactor facilities”, Transactions of the 3rd International Conference on StructuralMechanics in Reactor Technology, North-Holland Publishing Co., Amsterdam, 1975, 14pages, Vol. 4, Paper K 5/1.

Park, Y. J.; Wen, Y. K.; Ang, A. H.-S (1986). “Random vibration of hysteretic systemsunder bi-directional ground motions”, Earthquake Engineering & Structural Dynamics,14, 4, July-Aug. 1986, pages 543-557.

PEER (2000). “PEER Strong Motion Database”, processed by Dr. Walt Silva of PacificEngineering, Pacific Earthquake Engineering Research center, http://peer.berkeley.edu/scmat/, 2000.

SAC (1997). "Suites of Earthquake Ground Motions for Analysis of Steel Moment FrameStructures", Report No. SAC/BD-97/03, Woodward-Clyde Federal Services, SAC SteelProject.

Sheng, L. H., Hwang, J. S.; Gates, J. H (1994). “A comparison of equivalent elasticanalysis methods of bridge isolation”, Fifth U.S. National Conference on EarthquakeEngineering, Proceedings, Earthquake Engineering Research Inst., Oakland, California,Vol. I, 1994, pages 891-900.

Shibata, A., and Sozen, M.A. (1976). "Substitute structure method for seismic design in R/C", Journal of the Structural Division, ASCE, 102, ST1, pp1-18.

425

Somerville, P. (1997). “Engineering characteristicts of near fault ground motion”,SMIP97 Seminar on Utilization of Strong-Motion Data: Los Angeles, California, May 8,1997: Proceedings, California Division of Mines and Geology, Sacramento, 1997, pages9-28.

Stanton, J. F.; Roeder, C. W. (1991). “Advantages and limitations of seismic isolation”,400/E2395 Earthquake Spectra, 7, 2, May 1991, pages 301-324.

Stanton, John F. (1993). “Design criteria for seismic connections in precast concrete”,Structural Engineering in Natural Hazards Mitigation: Proceedings of Papers Presented atthe Structures Congress '93, American Society of Civil Engineers, New York, 1993, pages71-76, Vol. 1.

Stojadinovic, B.; Mackie, K.; (2002). “Design parameter sensitivity in bridge probabilisticseismic demand models”, Seventh U.S. National Conference on Earthquake Engineering(7NCEE), Earthquake Engineering Research Institute, Oakland, California.

Tung, T. P.; Newmark, N. M. (1954). “A review of numerical integration methods fordynamic response of structures”, University of Illinois at Urbana-Champaign., Urbana,Ill., 1954, 65, 18 pages.

U.B.C. (1994). “Division III - Earthquake Regulations for Seismic-Isolated Structures”,Appendix Chapter 16, Uniform Building Code, International Conference of BuildingOfficials, 1994.

U.S.D.O.T. (1997). “Seismic design of bridges, design example no. 2, 7, and 5“, UnitedStates Department of Transportation, Federal Highway Administration, Publication No.FHWA-SA-97-007, -010, and -012, October 1996.

Wallace, J.W. (1995). "Displacement-based design of R/C structures walls", Proceedingsof the 10th European Conference on Earthquake Engineering, 3, pp1539-144.

Whittaker, Andrew S. and Constantinou, Michael C. (1998). “Seismic protective systemsfor bridges in the United States: state-of-the-art and state-of-practice”, Nov. 3, 1998MCEER-98-0015, Proceedings of the U.S.-Italy Workshop on Seismic Protective Systemsfor Bridges, Multidisciplinary Center for Earthquake Engineering Research, Buffalo,N.Y., Nov. 3, 1998, pages 13-27.

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Performance-Based Design of Seismically Isolated Bridges A