Large Scale Earthquake 

Simulation of a Hybrid Lead 

Rubber Isolation System 

Designed with Consideration 

of Nuclear Seismicity

Office of Nuclear Regulatory Research

NUREG/CR-7196

CCEER 13-09
AVAILABILITY OF REFERENCE MATERIALS

IN NRC PUBLICATIONS

NRC Reference Material

As of November 1999, you may electronically access 

NUREG-series publications and other NRC records at 

NRC’s Library at www.nrc.gov/reading-rm.html. Publicly 

released records include, to name a few, NUREG-series

publications; Federal Register notices; applicant, 

licensee, and vendor documents and correspondence; 

NRC correspondence and internal memoranda; bulletins 

and information notices; inspection and investigative 

reports; licensee event reports; and Commission papers 

and their attachments.

NRC publications in the NUREG series, NRC regulations, 

and Title 10, “Energy,” in the Code of Federal Regulations

may also be purchased from one of these two sources.

1. The Superintendent of Documents

U.S. Government Publishing Office

Mail Stop IDCC

Washington, DC 20402-0001

Internet: bookstore.gpo.gov

Telephone: (202) 512-1800

Fax: (202) 512-2104

2. The National Technical Information Service

5301 Shawnee Rd., Alexandria, VA 22312-0002

www.ntis.gov

1-800-553-6847 or, locally, (703) 605-6000

A single copy of each NRC draft report for comment is

available free, to the extent of supply, upon written

request as follows:

Address: U.S. Nuclear Regulatory Commission

Office of Administration

Publications Branch

Washington, DC 20555-0001

E-mail: distribution.resource@nrc.gov

Facsimile: (301) 415-2289

Some publications in the NUREG series that are posted 

at NRC’s Web site address www.nrc.gov/reading-rm/

doc-collections/nuregs are updated periodically and may 

differ from the last printed version. Although references to 

material found on a Web site bear the date the material 

was accessed, the material available on the date cited 

may subsequently be removed from the site.

Non-NRC Reference Material

Documents available from public and special technical 

libraries include all open literature items, such as books, 

journal articles, transactions, Federal Register notices, 

Federal and State legislation, and congressional reports. 

Such documents as theses, dissertations, foreign reports 

and translations, and non-NRC conference proceedings 

may be purchased from their sponsoring organization.

Copies of industry codes and standards used in a

substantive manner in the NRC regulatory process are 

maintained at—

The NRC Technical Library

Two White Flint North

11545 Rockville Pike

Rockville, MD 20852-2738

These standards are available in the library for reference 

use by the public. Codes and standards are usually 

copyrighted and may be purchased from the originating 

organization or, if they are American National Standards, 

from—

American National Standards Institute

11 West 42nd Street

New York, NY 10036-8002

www.ansi.org

(212) 642-4900

Legally binding regulatory requirements are stated only in 

laws; NRC regulations; licenses, including technical specifications; or orders, not in NUREG-series publications. The 

views expressed in contractorprepared publications in this 

series are not necessarily those of the NRC.

The NUREG series comprises (1) technical and administrative reports and books prepared by the staff (NUREG–

XXXX) or agency contractors (NUREG/CR–XXXX), (2) 

proceedings of conferences (NUREG/CP–XXXX), (3) reports 

resulting from international agreements (NUREG/IA–XXXX), 

(4) brochures (NUREG/BR–XXXX), and (5) compilations of 

legal decisions and orders of the Commission and Atomic 

and Safety Licensing Boards and of Directors’ decisions 

under Section 2.206 of NRC’s regulations (NUREG–0750).

DISCLAIMER: This report was prepared as an account 

of work sponsored by an agency of the U.S. Government. 

Neither the U.S. Government nor any agency thereof, nor 

any employee, makes any warranty, expressed or implied, 

or assumes any legal liability or responsibility for any third 

party’s use, or the results of such use, of any information, 

apparatus, product, or process disclosed in this publication, 

or represents that its use by such third party would not 

infringe privately owned rights.
Large Scale Earthquake 

Simulation of a Hybrid Lead 

Rubber Isolation System 

Designed with Consideration 

of Nuclear Seismicity

Manuscript Completed: April 2013

Date Published: August 2015

Prepared by:

K. L. Ryan, C. B. Coria, N. D. Dao

Center for Civil Engineering Earthquake Research

(CCEER) University of Nevada, Reno/MS 0258

Reno, Nevada 89557-0258

Hernando Candra, NRC Project Manager

Office of Nuclear Regulatory Research

NUREG/CR-7196

CCEER 13-09

iii

ABSTRACT

Seismic isolation systems have been proven to provide superior performance and meet

continued functionality performance objectives for many facilities around the world, and are 

thus being considered for the future generation of nuclear power plants in the United States.

Experimental simulation of a hybrid lead-rubber isolation system for a 5-story steel moment

frame was performed at Hyogo Earthquake Engineering Research Center (E-Defense) of the

National Institute for Earth Science and Disaster Prevention in Japan. The isolation system was 

developed for the seismicity of a potential nuclear site in Central and Eastern United States

(CEUS). The isolation system was tested to displacements representing beyond design basis 

ground motions at the CEUS site and design basis ground motions for a Western United

States. Forces in the lead- rubber (LR) bearings were measured by an assembly of load cells.

The design of the isolation system was constrained by the experimental setup. The light

axial loads on the system, which are not representative of a nuclear facility, necessitated the

use of a hybrid system of bearings and flat sliders, known as cross linear (CL) bearings. The 

CL bearings support beneath some of the columns without contributing to the system base

shear, so that the desired isolation period could be provided at the target displacement.

Additionally, the CL bearings provided substantial resistance against the tensile demands

generated by overturning as a result of the light axial loads. Thus, the suitability of a hybrid

isolation system for nuclear power plant was evaluated as part of the test program.

A numerical simulation model was developed for the isolation system and the structure. The 

lead-rubber bearings were modeled with a bilinear force-displacement relation. Due to the 

amplitude dependence of the bearing response, the parameters of the bilinear model were

calibrated independently for each simulation. Using the calibrated model, the predicted 

displacement demand of the isolators was within 10% of the observed experimental 

displacement.

For typical XY (horizontal only) input excitation, the horizontal accelerations at the roof level 

were reduced by a factor of 10 relative to the fixed-base building. Under 3D input excitation 

which included vertical shaking, the vertical accelerations of all floor slabs were amplified by a 

factor of 4-6 relative to the input vertical excitation, but the vertical amplification factor was 

essentially the same for the isolated and the fixed-base building; thus the isolation system did

not increase vertical acceleration relative to the fixed-base building. Horizontal-vertical 

coupling was detected in the vibration modes of the structure (both fixed-base and baseisolated). The coupling was attributed partially to a mass irregularity. The majority of the 

coupling effects were replicated by a well-crafted numerical simulation model that accounted for

slab-frame interaction and refined distribution of mass over the floor system. The design of 

the base isolation system and structure should consider and accommodate these predictable 

horizontal-vertical coupling effects.

v 

FOREWORD

This NUREG/CR-7196 was prepared by the University of Nevada in Reno through commercial 

contract NRC-HQ-11-C-04-0067 (JCN V6274) with the Seismic Geotechnical Structural 

Engineering Branch, Division of Engineering Office of Nuclear Regulatory Research. This 

research was part of a larger collaboration between the National Science Foundation (NSF) 

Network for Earthquake Engineering Simulation (NEES) and Japan’s National Research 

Institute for Earth Science and Disaster Prevention (NIED). This collaboration created the 

opportunity and a cost effective way to test and investigate Hybrid Lead Rubber Isolation System

isolation system focused on nuclear applications.

Through pre-application meetings between NRC staff and vendors, staff identified the body of 

research and development already done in the seismic isolation (SI) area. Based on those 

meetings NRC staff anticipates that the types of isolators investigated by this research are those 

that would be used in design applications. The data that this testing program was designed to 

acquire are anticipated to address a gap in the NRC staff’s ability to review applications that 

include the use of rubber bearing SI devices and the performance of these isolators under the 

extreme beyond-design-basis ground motions for the staff review of seismic safety margins. The 

resulting data produced by this research can also be used to verify the confirmatory numerical 

models for SI devices in the new NRC ESSI Simulator code, a seismic soil-structure interaction 

tool under development by the NRC and the University of California at Davis. 

This report provides the technical information and comprehensive test data that can be used by 

the NRC staff to support new regulatory guidance related to seismic isolation (SI) technology on 

future nuclear power plant structures. Despite the light weight and the application of cross linear

(CL) bearings on the test system (hybrid system), which are not representative of a nuclear 

facility, the CL bearings provided substantial resistance against the tensile demands and large 

displacement stability without contributing to the system base shear. Thus, these test data can 

be used for numerical confirmatory analysis and benchmarking for SI, and advance the steps 

toward the solution of key technical issues that must be addressed before base isolation 

systems can safely be implemented for the economical seismic protection of nuclear facilities.

A numerical simulation with OpenSees code was used for this isolation system and the 

structure. Due to OpenSees lacks of features compared with commercial software (Ansys, 

Abaqus, or LS Dyna), many assumptions and approximations were used in this analysis. The 

CL bearings were represented by a combination of spring elements due to lack of contact 

algorithm, this approximation is not strictly accurate since the CL bearing is a friction device 

which can slide in two directions, and its resistant is proportional to the instantaneous axial 

force. 

The lead-rubber bearings that were modeled with a bilinear force-displacement relation needs to 

be re-calibrated independently for each simulation to match the test data. This indicates the

need of improved models that can capture others nonlinear parameters such as the effects of 

heating of the lead core, pressure variation, and rate dependence (research on this device is 

under development by the NRC and the University of Buffalo). Because of the difficulties in 

modeling the response of the tested LR bearings, the simulation results should be carefully 

studied before using them in supporting the regulatory guidance, however the analysis still 

provides some useful insight and the test data can be used for confirmatory analysis and 

simulation benchmark.

vii

TABLE OF CONTENTS

ABSTRACT................................................................................................................................iii

FOREWORD .............................................................................................................................. v

TABLE OF CONTENTS ............................................................................................................ vii

LIST OF FIGURES ....................................................................................................................xi

LIST OF TABLES..................................................................................................................... xxi

EXECUTIVE SUMMARY.........................................................................................................xxiii

ACKNOWLEDGMENT............................................................................................................ xxv

NOMENCLATURES............................................................................................................... xxvii 

1. INTRODUCTION................................................................................................................. 1

1.1 Potential for Seismic Isolation of Safety Related Nuclear Facilities.......................... 1

1.2 State of Knowledge and Motivation for Full Scale Testing ....................................... 3

1.3 Opportunity for Testing at E-Defense ...................................................................... 4

1.4 Report Organization ................................................................................................ 7 

2. DESIGN AND CONFIGURATION OF THE 5-STORY STEEL MOMENT FRAME

TESTBED STRUCTURE.................................................................................................... 9

2.1 Description of Testbed Structure ............................................................................. 9

2.1.1 Basic Dimensions.................................................................................................... 9

2.1.2 Design Spectra and Design Criteria...............................................................10

2.1.3 Framing System ............................................................................................10

2.1.4 Slabs .............................................................................................................13

2.1.5 Material Properties ........................................................................................13

2.2 Non-Structural Components and Contents .............................................................14

2.3 Weights..................................................................................................................15

2.4 System Identification ..............................................................................................17

2.5 Condition of the Testbed Prior to Testing ...............................................................18

3. DESIGN OF THE ISOLATION SYSTEM............................................................................19

3.1 Target Response of a Nuclear Site.........................................................................19

3.2 Isolation System Design.........................................................................................21

3.2.1 Estimated Displacement Demands..............................................................21

3.2.2 Configuration of Lead-Rubber Bearings.......................................................22

3.2.3 Selection of the Bearing Dimensions ...........................................................26

3.2.4 Selection of the Lead Plug Dimensions .......................................................27

3.2.5 Cross Linear Bearings .................................................................................28

3.2.6 Summary of Design Properties....................................................................30

3.3 Selection and Scaling of Ground Motions...............................................................32
viii

3.3.1 Imposed Limitations for Safety ......................................................................33

3.3.2 Vogtle and Diablo Canyon Motions................................................................34

3.3.3 Remainder of the Test Program ....................................................................38

4. EXPERIMENTAL PROGRAM ...........................................................................................41

4.1 Design of Connection Assembly .............................................................................41

4.2 Instrumentation ......................................................................................................44

4.2.1 Load Cells .....................................................................................................45

4.2.2 Displacement Transducers............................................................................45

4.2.3 Accelerometers .............................................................................................48

4.3 Installation of the Specimen on the Earthquake Simulator......................................50

4.4 Test Schedule ........................................................................................................54

4.5 Table Motions.........................................................................................................58

4.6 Derived Responses................................................................................................64

4.6.1 Horizontal Displacement of the Isolation System..........................................64

4.6.2 Isolator Forces..............................................................................................67

4.6.3 Horizontal Acceleration and Story Drifts........................................................68

5. SUMMARY OF EXPERIMENTAL SIMULATION RESULTS FOR THE HYBRID LR

ISOLATED BUILDING.......................................................................................................71

5.1 Isolator Displacements ...........................................................................................71

5.2 Torsional Response ...............................................................................................74

5.3 System Base Shear................................................................................................75

5.4 Axial Forces in LR Bearings ...................................................................................78

5.5 Isolation System Re-Centering...............................................................................81

5.6 Floor Accelerations in the Testbed Building............................................................84

5.7 Story Drifts in Testbed Building ..............................................................................87

6. TECHNICAL RESPONSE OF HYBRID LR ISOLATION SYSTEM.....................................89

6.1 Bolt Slip in LR Bearings..........................................................................................89

6.2 Transfer of Load Between Bearings .......................................................................96

6.3 Repeatability of Isolation System Response after Many Tests .............................108

7. MODEL FOR ISOLATION BEARINGS AND BEARING CHARACTERIZATION ..............121

7.1 Lead-Rubber Bearings .........................................................................................121

7.1.1 Horizontal Direction Modeling Assumptions................................................121

7.1.2 Vertical Direction Modeling Assumptions ................................................... 122

7.2 Cross Linear Bearing............................................................................................123

7.2.1 Horizontal Direction Modeling Assumptions................................................123

7.2.2 Vertical Direction ........................................................................................124

7.3 Composite Vertical Properties of the Isolation System ........................................125

7.4 Characterization of Lead-Rubber Bearings..........................................................125
ix

7.4.1 Characterization by Dynamic Isolation Systems (DIS)................................126

7.4.2 Characterization Based on Experimental Data ...........................................129

7.4.3 Trilinear Characterization ...........................................................................136

7.4.4 Final Bearing Parameter Selection for Numerical Simulation ......................139

8. NUMERICAL SIMULATION MODEL OF TESTBED STRUCTURE................................. 141

8.1 Modeling Beams and Columns.............................................................................142

8.2 Modeling Panel Zones..........................................................................................144

8.3 Modeling Gravity Load and Mass .........................................................................146

8.4 Modeling Damping ...............................................................................................147

9. COMPARISON OF EXPERIMENTAL AND NUMERICAL SIMULATION .........................151

9.1 Isolator Response ................................................................................................151

9.2 Modal Analysis .....................................................................................................161

9.3 Structural Response.............................................................................................168

10. INFLUENCE OF VERTICAL EXCITATION ON THE STRUCTURAL RESPONSE ..........181

10.1 Response of LR Bearings under Axial Force Variation.........................................181

10.2 Horizontal and Vertical Floor System Vibration.....................................................185

10.3 Source of Amplification of Horizontal Floor Acceleration under 3D Input ..............189

10.4 Influence of Supplementary Roof Mass on Vertical Slab Vibration and 

Horizontal-Vertical Coupling.................................................................................195

10.5 Implications for Design.........................................................................................201

11. STRUCTURAL AND ISOLATOR RESPONSE FOR THREE BUILDING

CONFIGURATIONS........................................................................................................203

11.1 Structural Response.............................................................................................205

11.2 Isolation System Response..................................................................................211

12. CONCLUSIONS AND CONSIDERATIONS FOR NUCLEAR POWER PLANTS ...............221

12.1 Characteristics of Lead-Rubber Bearing Response..............................................221

12.2 Predictability of the System Response ................................................................. 221

12.3 Hybrid Lead-Rubber (LR) and Cross-Linear (CL) Bearing System .......................223

12.4 Comparison of the Three Building Configurations ................................................224

12.5 Influence of Vertical Excitation on the Structural Response..................................225

13. REFERENCES .................................................................................................................227
x 

APPENDIX A DESIGN & CONSTRUCTION DRAWING FOR TESTING OF VALUEADDED DAMPED BUILDING ....................................................................... A-1 

APPENDIX B DRAWINGS AND SPECIFICATIONS FOR LEAD-RUBBER BEARINGS 

AND CROSS LINEAR BEARINGS................................................................ B-1 

APPENDIX C DRAWINGS FOR LOAD CELL ASSEMBLIES AND CONNECTION TO 

THE SIMULATOR PLATFORM .................................................................... C-1 

APPENDIX D DRAWINGS FOR STRUCTURAL INSTRUMENTATION PLAN.................... D-1 

APPENDIX E ISOLATOR TEST REPORT FOR LEAD-RUBBER BEARINGS .................E-1 
xi

LIST OF FIGURES 

Figure 1-1 E-Defense earthquake simulator parameters ..................................................... 5

Figure 2-1 The 5-story steel moment frame specimen ......................................................... 9

Figure 2-2 Basic dimensions of the specimen: (a) typical plan view from floors 2 to 5, 

and (b) elevation view ......................................................................................10

Figure 2-3 Design spectrum for Japanese Level II and Level III earthquakes ....................11

Figure 2-4 Beam, beam to column connection and slab ....................................................11

Figure 2-5 Column base: (a) view from top (b) view from bottom, (c) rendering of 

stiffeners ..........................................................................................................12

Figure 2-6 Horizontal braces at base level .........................................................................13

Figure 2-7 Hospital room ...................................................................................................15

Figure 2-8 Office room. ......................................................................................................15

Figure 2-9 Location of steel weights at roof .......................................................................16

Figure 2-10 Transfer function during white noise excitation in fixed-base configuration: 

(a) x-direction and (b) y-direction .....................................................................17

Figure 2-11 Cracks in concrete slab and rust on steel member. ...........................................18

Figure 3-1 5% damped URS for the DBE at the Vogtle site. (Source: Huang et 

al.2009)............... ............................................................................................20

Figure 3-2 5% damped response spectra for the 30 sets of spectrum compatible 

maximum-minimum motions: (a) maximum component, (b) minimum 

component and (c) vertical component. (Source: Huang et al. 2009) .................20

Figure 3-3 Idealized bilinear force-displacement relationship for LR bearings, 

determined by post-yield stiffness Kd (stiffness of rubber), characteristic 

strength Qd (strength of lead plug) and initial stiffness K1 or yield 

displacement Dy. An effective or secant stiffness Keff is determined 

as the peak-to-peak stiffness based on maximum force Fmax and 

displacement Dmax .......................... ..........................................................................21

Figure 3-4 Displacement demand representing (a) median response to maximumminimum motions and (b) 90% percentile response to 150% maximumminimum motions for Vogtle site .......... ............................................................22

Figure 3-5 Drawing of testbed base plan with column labels (N = North, E = East, 

S = South, W = West, C = Center) .....................................................................23

Figure 3-6 Illustration of a connection assembly with triaxial load cells to measure 

bearing forces ....................................................................................................23

Figure 3-7 (a) Configuration Option 1 with LR bearings beneath 4 corner columns, 

(b) Configuration Option 2 with LR bearings (circles) beneath 4 edge 

columns and flat sliders (squares) beneath remaining columns. The 

supported weights (in units of kN) at the isolators based on tributary 

loads are also indicated .....................................................................................25

Figure 3-8 Proposed design options for LR bearings ..........................................................26

Figure 3-9 Axial force capacity of proposed LR bearings versus lateral displacement ........27

Figure 3-10 Photo of installed CL bearing illustrating orthogonal LM guides assemblies 

on top and bottom ..............................................................................................29
xii 

Figure 3-11 Total base shear (Vb) and overturning forces (FOT) acting on an elevation 

view of the testbed building ................................................................................30

Figure 3-12 (a) Final plan drawing of the hybrid LR isolation system, (b) photo of 

LR bearing, and (c) photo of CL bearing ............................................................31

Figure 3-13 Calculated displacement trace of the isolation system for 6 Vogtle 

motions by SDOF analysis ...... ...............................................................34

Figure 3-14 Calculated displacement histories of the isolation system in x and ydirections for 6 Vogtle motions by SDOF analysis ..............................................35

Figure 3-15 Predicted displacement trace of each isolator for the OpenSees building 

model subjected to 150% Vogtle #13 ................................................................. 36

Figure 3-16 Predicted displacement trace of each isolator for the OpenSees building 

model subjected to 150% Vogtle #13, with input ground excitation 

components rotated by 11.25 degrees ...... ................................................37

Figure 3-17 Predicted displacement trace of each isolator for the OpenSees building 

model subjected to 100% Diablo Canyon #15 ....................................................38

Figure 4-1 Load cell connection assemblies used for: (a) East bearing and (b) North, 

South and West bearings ...................................................................................42

Figure 4-2 Connection assembly ........................................................................................42

Figure 4-3 Load cell drawings with bolt patterns and elevation views: (a) Type A and 

(b) Type B ..........................................................................................................43

Figure 4-4 Connection of Type A and Type B load cells to top and bottom connection 

plates .................................................................................................................44

Figure 4-5 Magnitude of low-pass Butterworth Filter transfer function ................................. 45

Figure 4-6 Load cell channels for the hybrid LR isolation configuration ...............................46

Figure 4-7 Layout of displacement transducers at base ......................................................47

Figure 4-8 Instrumentation for measuring story drift ............................................................47

Figure 4-9 Layout of displacement transducers to measure story drift in 2nd to 5th

stories ........................ ...................................................................................48

Figure 4-10 Accelerometers at the top connection plates to estimate inertia forces ..............49

Figure 4-11 Layout of accelerometers at the 5th floor ............................................................49

Figure 4-12 Connection assembly .........................................................................................50

Figure 4-13 Connection assemblies on the simulator platform ..............................................51

Figure 4-14 Installing the LR bearings to the connecting assemblies and the CL 

bearings to the simulator platform ......................................................................51

Figure 4-15 Lowering the testbed building over the isolation system ....................................52

Figure 4-16 Bolt holes for connecting the testbed structure to the isolation system ..............52

Figure 4-17 Drilling and tapping holes at the bottom of the specimen ...................................53

Figure 4-18 Acceleration history of target and realized RRS88 motion in the hybrid LR 

isolation configuration test ..................................................................................61

Figure 4-19 Ratio of realized to target motion 5% damped spectral accelerations – 

hybrid LR configuration ......................................................................................62

Figure 4-20 Ratio of realized to target motion 5% damped spectral accelerations – 

fixed-based configuration ...................................................................................63

Figure 4-21 Configurations for solving displacement of the isolation system. (a) 

original configuration, (b) displaced configuration...............................................66

Figure 4-22 Free body diagram illustrating derivation of isolator reaction .............................67
xiii 

Figure 4-23 Diagram illustrating the computation of drift at the geometric center ...................69

Figure 5-1 Peak vector sum displacement recorded in any LR bearing for each 

earthquake simulation ........... ..........................................................................72

Figure 5-2 X-direction, y-direction and overall peak (vector sum) displacement 

of each LR bearing for each earthquake simulation ...........................................72

Figure 5-3 Displacement trace (x vs y-direction displacement) of the East LRB for 

(a) WSM80, (b) DIA95(XY), (c) VOG100 and (d) VOG175 ................................. 73

Figure 5-4 Peak rotation angle of the base for each earthquake simulation ........................74

Figure 5-5 Displacement trace of each isolator during the Vogtle 175% simulation ............75

Figure 5-6 Base shear coefficient for all simulations: total, x and y-directions .....................77

Figure 5-7 Base shear coefficient for each simulation alongside the design backbone 

curve ..................................................................................................................78

Figure 5-8 Peak compressive force in any LR bearing for each earthquake simulation .......80

Figure 5-9 Peak tensile force in any LR bearing for each earthquake simulation. 

(A tensile force of zero indicates that tension was not observed) ......................80

Figure 5-10 Peak axial forces in each LR bearing for each simulation: (a) Peak 

compression force, (b) peak normalized compression force, (c) peak 

tension force, and (d) normalized peak tension force .........................................81

Figure 5-11 X-direction, y-direction, and total (vector sum) displacement recorded 

in each LR bearing at the end of every earthquake simulation ...........................82

Figure 5-12 Permanent displacement of around 1.1 cm on the East bearing due to 

sliding of the bottom steel plate ..........................................................................83

Figure 5-13 Permanent displacement at the location of each LR bearing at the end 

of one simulation compared to the beginning of the next....................................84

Figure 5-14 Peak acceleration profile for all simulations in both horizontal directions ............86

Figure 5-15 Peak floor acceleration profiles for increasing intensity of Vogtle input 

excitation (75% - 175%) ...... .............................................................................86

Figure 5-16 Peak absolute and normalized acceleration profile comparing XY and 

3D excitations for Diablo (95% and 80% respectively) and Rinaldi (88%) ..........87

Figure 5-17 Peak story drift throughout the height of the building for all excitations in 

x and y directions ...............................................................................................88

Figure 6-1 Movement of (a) East LR bearing relative to (b) top and (c) bottom plates. ........90

Figure 6-2 Horizontal force history of the East bearing (LRB-E) in the x- and ydirections for a subset of the trials ......................................................................91

Figure 6-3 X and y-direction hysteresis loops (horizontal force vs. displacement) 

of the East bearing (LRB-E) during Vogtle 125% and Vogtle 150% ...................93

Figure 6-4 Horizontal (Fx and Fy) and vertical (Fz) force history of all four LR beaings 

during Vogtle 150%......... ...................................................................................94

Figure 6-5 Propagation of bolt slip through the height of the structure in x-direction 

during Vogtle 150%; unfiltered horizontal force in LRB-E and 1st – 6th floor 

acceleration in SE column...... ...........................................................................95

Figure 6-6 Resultant action on LR bearings as a result of CL bearings and base 

diaphragm constraint........... .............................................................................96

Figure 6-7 History of average horizontal displacement (x, y and vector sum),and 

axial force in individual LR bearings and summed over all LR bearings 

for Sine Wave (XY) ............................................................................................97
xiv

Figure 6-8 History of average horizontal displacement (x, y and vector sum), and 

axial force in individual LR bearings and summed over all LR bearings 

for Diablo 95% (XY) ...........................................................................................98

Figure 6-9 History of average horizontal displacement (x, y and vector sum), and 

axial force in individual LR bearings and summed over all LR bearings for 

Rinaldi 88% (XY)................................................................................................99

Figure 6-10 History of axial force in individual LR bearings and summed over all 

LR bearings for Rinaldi 88% .................................................................101

Figure 6-11 History of axial force in individual LR bearings and summed over all LR 

bearings comparing Rinaldi 88% – dashed line – and Rinaldi 88% (XY) – 

solid line. A low pass filter with a cutoff frequency of 2 Hz was applied to 

the total axial force for Rinaldi 88% and is superimposed over the total, 

shown as a red dashed line .............................................................................102

Figure 6-12 History of average vector sum displacement at building center, axial force 

in individual LR bearings and axial force summed over all LR bearings 

comparing Diablo 95% – dashed line – and Diablo 80% (XY) – solid line. 

A low pass filter with a cutoff frequency of 2 Hz was applied to the total 

axial force for Diablo 80% and is superimposed over the total, shown 

as a red dashed line .........................................................................................103

Figure 6-13 History of average vector sum displacement at building center, axial force 

in individual LR bearings and axial force summed over all LR bearings for 

Vogtle 175%. A low pass filter with a cutoff frequency of 2 Hz was applied 

to the total axial force and is superimposed over the total, shown as a red 

dashed line ......................................................................................................104

Figure 6-14 (a) Relative change in compressive force and (b) Total compressive force 

summed over LR bearings at the start and end of all simulations .....................106

Figure 6-15 Sample feedback acceleration at the center of the earthquake simulator 

(averaged over several accelerometers) in the x and y-directions 

compared for the first and second Sine 100% (Y) and Vogtle 75% 

simulations ......... ...........................................................................................109

Figure 6-16 5% damped response spectra for the feedback acceleration at the center 

of the earthquake simulator (averaged over several accelerometers) in the 

x, y and z-directions compared for the first and second Sine 100% (Y) 

and Vogtle 75% simulations ............. .............................................................110

Figure 6-17

Figure 6-18

Y-direction force vs. displacement (hysteresis loops) for the 1st and 2nd 

simulation of Sine 100%(Y) ..................................................................112

Y-direction displacement and force histories for the 1st and 2nd simulation 

of Sine 100% (Y) ..............................................................................................113

Figure 6-19 X-direction force vs. displacement (hysteresis loops) for the 1st and 2nd

simulation of Vogtle 75% ..................................................................................114

Figure 6-20 Y-direction force vs. displacement (hysteresis loops) for the 1st and 2nd

simulation of Vogtle 75% ..................................................................................115

Figure 6-21 X-direction displacement and force histories for the 1st and 2nd simulation 

of Vogtle 75% ..................................................................................................116

Figure 6-22 Y-direction displacement and force histories for the 1st and 2nd simulation 

of Vogtle 75% ..................................................................................................117
xv 

Figure 6-23 X vs. y-direction displacement (displacement trace) for the 1st and 2nd

simulation of Vogtle 75% ..... .................................................................118

Figure 6-24 Y-direction force vs. displacement (hysteresis loops) identified by cycle 

for the 1st and 2nd simulation of Sine 100% (Y) .................................................119

Figure 7-1 Horizontal force-displacement of LR bearing for numerical simulation ..............122

Figure 7-2 Vertical force-deformation of LR bearing for analytical model: (a) 

elastic-no tension, (b) elastic, and (c) combined ..............................................122

Figure 7-3 Horizontal force-displacement of CL bearing for numerical simulation ..............124

Figure 7-4 Vertical force-deformation of CL bearing for analytical model: (a) elastic-no 

tension, (b) elastic, and (c) combined ...............................................................125

Figure 7-5 Comparison of fitted hysteresis loop and parameters to test data .....................127

Figure 7-6 Hysteresis loop of Test A and D for one of the isolation pairs ...........................128

Figure 7-7 Projection of displacement trace to main axis for Vogtle 100% .........................129

Figure 7-8 Selection of cycle for characterization for Vogtle 100% ....................................130

Figure 7-9 Hysteresis loop of peak cycle for the projected-direction for East 

bearing for Vogtle 100%...................................................................................131

Figure 7-10 Analytical and experimental hysteresis loop using (a),(c) isolator design 

parameters and (b),(d) characterized parameters ............................................134

Figure 7-11 Y-direction hysteresis and force history of test data of the East bearing 

compared to (a - b) characterization model,and (c - d) design model for 

Sine 100%(Y)-1. ..............................................................................................135

Figure 7-12 X and y-direction displacement history of experimental data compared to 

(a) characterized model and (b) design model for Vogtle 175%, East 

bearing ............. .............................................................................................136

Figure 7-13 Force vs. displacement loop for the East LRB in Sine 100%(Y)-1 ....................137

Figure 7-14 Conceptual force vs. displacement of lead plug for monotonic loading 

(courtesy of Dynamic Isolation Systems) .........................................................137

Figure 7-15 Comparison of numerical and experimental simulation force versus 

displacement of the East bearing for SIN100(Y)-1 and VOG100 using (a), 

(c) a trilinear hysteretic model and (b), (d) a bilinear hysteretic model ..............139

Figure 7-16 Comparison of numerical simulation with 100% and 85% characterized 

values of Qd to the experimental data for Vogtle 100%; displacement 

history and hysteresis loop in the y-direction for the East bearing ....................140

Figure 8-1 (a) Photo and (b) OpenSees model of the testbed building without isolators ....141

Figure 8-2 Primary beam, beam-to-column connection, and slab .....................................142

Figure 8-3 Behavior of a representative composite fiber section beam with and 

without axial restraint: (a) fiber section geometry and (b) forcedisplacement relationship .................................................................................143

Figure 8-4 Composite section behavior: (a) moment-curvature relationship of the 

section, (b) component 1 of resultant section modeled by steel material 

model, and (c) component 2 of resultant section modeled by hysteresis 

model ...............................................................................................................144

Figure 8-5 Panel zone model for beam to column connection. (a) beam to column 

connection, (b) numerical model of panel zone ................................................145

Figure 8-6 Gusset plate and its finite element model: (a) gusset plate, (b) Von-Mises 

stress due to gravity load .................................................................................146
xvi

Figure 8-7 Rayleigh damping model for the fixed-base and hybrid LR isolated 

building model ..................................................................................................148

Figure 8-8 Additional damper for adjusting the damping coefficient of the 2nd mode 

in the X direction ..............................................................................................149

Figure 9-1 Comparison of numerical and experimental simulation for El Centro 

130%; displacement trace and displacement histories (x and y) in each LR 

bearing, labeled by position (E, S, N, W). Numerical simulation uses the 

characterized model with 85% of the characterized value for Qd ...................153

Figure 9-2 Comparison of numerical and experimental simulation for El Centro 130%; 

bearing force vs. displacement or hysteresis loops (x and y) in each LR 

bearing, labeled by position (E, S, N, W). Numerical simulation uses the 

characterized model with 85% of the characterized value for Qd ......................154

Figure 9-3 Comparison of numerical and experimental simulation for Vogtle 100%; 

displacement trace and displacement histories (x and y) in each LR 

bearing, labeled by position (E, S, N, W). Numerical simulation uses the 

characterized model with 85% of the characterized value for Qd ......................155

Figure 9-4 Comparison displacemet (x and y) histories of numerical and experimental 

simulation for Vogtle 100% in each LR bearing labeled my position 

(E, S, N, W). Numerical simulation uses the characterized model with 85% 

of the characterized value for Qd ......................................................................156

Figure 9-5 Comparison of numerical and experimental simulation for Vogtle 175%; 

displacement trace and displacement histories (x and y) in each LR 

bearing, labeled by position (E, S, N, W). Numerical simulation uses the 

characterized model with 85% of the characterized value for Qd ......................157

Figure 9-6 Comparison of numerical and experimental simulation for Vogtle 175%; 

displacement trace and displacement histories (x and y) in each LR 

bearing, labeled by position (E, S, N, W). Numerical simulation uses the 

characterized model with 85% of the characterized value for Qd ......................158

Figure 9-7 Comparison of numerical and experimental simulation for Diablo 95%; 

displacement trace and displacement histories (x and y) in each LR 

bearing, labeled by position (E, S, N, W). Numerical simulation uses the 

characterized model with 85% of the characterized value for Qd ......................159

Figure 9-8 Comparison displacement (x and y) histories of numerical and experimental 

simulation for Diablo 95% in each LR bearing, labeled by position (E, S, N, 

W). Numerical simulation uses the characterized model with 85% of the 

characterized value for Qd................................................................................160

Figure 9-9 Comparison of numerical and experimental simulation for Vogtle 175%; 

history of base rotation angle. Analysis uses the characterized model with 

85% of the characterized value for Qd ..............................................................161

Figure 9-10 Modes 1-4 deflected shape of all the frames in 3-dimensional, x- and 

y-direction, along with their associated frequencies and periods. The 

undeflected shape of the system is shown in dashed lines ..............................163

Figure 9-11 Modes 5-8 deflected shape of all the frames in 3-dimensional, x- and 

y-direction, along with their associated frequencies and periods. The 

undeflected shape of the system is shown in dashed lines ..............................164
xvii

Figure 9-12 Modes 9-12 deflected shape of all the frames in 3-dimensional, x- and ydirection, along with their associated frequencies and periods. The 

undeflected shape of the system is shown in dashed lines. .............................165

Figure 9-13 Modes 13-16 deflected shape of all the frames in 3-dimensional, x- and ydirection, along with their associated frequencies and periods. The 

undeflected shape of the system is shown in dashed lines ..............................166

Figure 9-14 Modes 17-20 deflected shape of all the frames in 3-dimensional, x- and ydirection, along with their associated frequencies and periods. The 

undeflected shape of the system is shown in dashed lines ..............................167

Figure 9-15 Comparison of numerical and experimental simulation for El Centro 130%; 

x and y-direction roof acceleration histories, peak acceleration profiles, 

2nd story drift histories, and peak drift profiles. Numerical simulation 

uses the characterized model with 85% of the characterized value for Qd .......169

Figure 9-16 Comparison of numerical and experimental simulation for Vogtle 100%; 

x and y-direction roof acceleration histories, peak acceleration profiles, 

2nd story drift histories, and peak drift profiles. Numerical simulation uses 

the characterized model with 85% of the characterized value for Qd ................170

Figure 9-17 Comparison of numerical and experimental simulation for Vogtle 175%; 

x and y-direction roof acceleration histories, peak acceleration profiles, 

2nd story drift histories, and peak drift profiles. Numerical simulation 

uses the characterized model with 85% of the characterized value for Qd........171

Figure 9-18 Comparison of numerical and experimental simulation for Diablo 95%; 

x and y-direction roof acceleration histories, peak acceleration profiles, 

2nd story drift histories, and peak drift profiles. Numerical simulation uses 

the characterized model with 85% of the characterized value for Qd ................172

Figure 9-19 Comparison of numerical and experimental simulation for El Centro 130%; 

x and y-direction spectral accelerations for 1st through 6th floors. 

Numerical simulation uses the characterized model with 85% of the 

characterized value for Qd................................................................................174

Figure 9-20 Comparison of numerical and experimental simulation for Vogtle 100%; 

x and y-direction spectral accelerations for 1st through 6th floors. 

Numerical simulation uses the characterized model with 85% of the 

characterized value for Qd. ................................................................................175

Figure 9-21 Comparison of numerical and experimental simulation for Vogtle 175%; 

x and y-direction spectral accelerations for 1st through 6th floors. Numerical 

simulation uses the characterized model with 85% of the characterized 

value for Qd ......................................................................................................176

Figure 9-22 Comparison of numerical and experimental simulation for Diablo 95% (XY); 

x and y-direction spectral accelerations for 1st through 6th floors. Numerical 

simulation uses the characterized model with 85% of the characterized 

value for Qd ......................................................................................................177

Figure 10-1 Bearing force vs. displacement (hysteresis loops) in x-direction for a single 

large displacement cycle compared for Rinaldi 88% (XY) and Rinaldi 88% 

(V-PGA = 1.21g) excitations ...........................................................................182
xviii

Figure 10-2 Bearing force vs. displacement (hysteresis loops) in y-direction for a 

single large displacement cycle compared for Rinaldi 88% (XY) and 

Rinaldi 88% (V-PGA = 1.21g) excitations .......................................................182

Figure 10-3 History of shear force and axial force variation in LRB-S compared for 

Rinaldi 88% (XY) and Rinaldi 88% (V-PGA = 1.21g) excitations ....................183

Figure 10-4 Bearing force vs. displacement (hysteresis loops) in x-direction for a 

single cycle compared for Diablo 80% (V-PGA = 0.44g) and Diablo 95% 

(XY) excitations. ............................................................................................ .184

Figure 10-5 Bearing force vs. displacement (hysteresis loops) in y-direction for a 

single cycle compared for Diablo 80% (V-PGA = 0.44g) and Diablo 95% 

(XY) excitations ............................................................................................. .184

Figure 10-6 History of shear force and axial force variation in LRB-N compared for 

Diablo 80% (V-PGA = 0.44g) and Diablo 95% (XY) excitations .....................185

Figure 10-7 Comparison of vertical roof slab acceleration (blue), 5th floor slab 

acceleration (red), and ground acceleration (black) for (a) Vogtle 100%, 

(b) Diablo 80%, and (c) Rinaldi 88%. Peak values indicated on graph ...........186

Figure 10-8 Comparison of numerical simulation to experimental data for vertical slab 

acceleration; (a) roof floor and (c) 5th floor for Diablo 80%, and (b) roof 

floor and (d) 5th floor for Rinaldi 88% ..............................................................187

Figure 10-9 Comparison of x and y-direction roof acceleration histories and peak 

acceleration profiles for Diablo 80% and Diablo 95% (XY) .............................188

Figure 10-10 Comparison of x and y-direction roof acceleration histories and peak 

acceleration profiles for Rinaldi 88% (XY) and Rinaldi 88%............................189

Figure 10-11 Comparison of x and y-direction 5% damped spectral accelerations for 1st

through 6th floors for Diablo 80% and Diablo 95% (XY). These spectra 

were derived from the experimentally recorded signals ..................................190

Figure 10-12 Comparison of x and y-direction 5% damped spectral accelerations for 1st

through 6th floors for Rinaldi 88% (XY) and Rinaldi 88%. These spectra 

were derived from the experimentally recorded signals ..................................191

Figure 10-13 Comparison of numerical simulation to experimental data for Diablo 80%; 

x and y-direction 5% damped spectral accelerations for 1st through 

6th floors. Numerical simulation uses the characterized bearing model 

with 85% of the characterized value for Qd.....................................................193

Figure 10-14 Comparison of numerical simulation to experimental data for Rinaldi 88%; 

x and y-direction 5% damped spectral accelerations for 1st through 6th 

floors. Numerical simulation uses the characterized bearing model with 

85% of the characterized value for Qd. ...........................................................194

Figure 10-15 Numerically simulated vertical slab acceleration for Diablo 80% and 

Rinaldi 88%; comparison using the testbed building model and the 

augmented model with supplementary mass removed from the roof 

and replaced at the base. ................................................................. ..............196 

Figure 10-16 Comparison of numerically simulated x and y-direction roof acceleration 

histories and peak acceleration profiles for Diablo 80% and Diablo 95% 

(XY) using the augmented model with supplementary mass removed 

from the roof and replaced at the base ...........................................................197
xix

Figure 10-17 Comparison of numerically simulated x and y-direction roof acceleration 

histories and peak acceleration profiles for Rinaldi 88% XY and Rinaldi 

88% using the augmented model with supplementary mass removed 

from the roof and replaced at the base ...........................................................198

Figure 10-18 Comparison of numerically simulated x and y-direction 5% damped 

spectral accelerations in 1st through 6th floors for Diablo 80% and Diablo 

95% (XY) using the augmented model with supplementary mass 

removed from the roof and replaced at the base ............................................199

Figure 10-19 Comparison of numerically simulated x and y-direction 5% damped 

spectral accelerations in 1st through 6th floors for Rinaldi 88% XY and 3D 

using the augmented model with supplementary mass removed from 

the roof and replaced at the base ...................................................................200

Figure 11-1 Normalized backbone curves (base shear vs. displacement) for the Long 

Period TP isolation system and the Moderate Period LR isolation system, 

based on design properties of the bearings and measured mass ...................203

Figure 11-2 Peak floor acceleration (vector sum) vs. floor level of the three building 

configurations for (a) WSM 80%, (b) IWA 100%/70%, (c) Rinaldi 

88%/35% (XY) and (d) Rinaldi 88%/35% records. Accelerations are 

absolute on the left graph and normalized by PGA on the right. .....................206

Figure 11-3 History of total base shear in x and y-directions compared for the 3 system 

configurations in (a) Rinaldi 88% (XY) and (b) Rinaldi 88%. The bottom 

graphs provide a close-up of the isolated configurations; peak values are 

indicated ........................................................................................................208

Figure 11-4 Comparison of x and y-direction 5% damped spectral accelerations in the 

TP isolation configuration for 1st through 6th floors for Rinaldi 88% XY 

and Rinaldi 88% .............................................................................................209

Figure 11-5 Vertical acceleration history of floor slab (recorded at middle of the NE 

quadrant), column average, and ground acceleration average 

(recorded at simulator platform) for Rinaldi 88% for (a,d) TP isolation 

configuration, (b,e) hybrid LR isolation configuration, and (c,f) fixed-base 

configuration; (a-c) = 5th floor slab and (d-f) = roof slab ..................................210

Figure 11-6 Peak story drift (max in x or y-direction) vs. story level of the three 

building configurations for (a) WSM 80%, (b) IWA 100%/70%, (c) Rinaldi 

88%/35% (XY) and (d) Rinaldi 88%/35% records ...........................................211

Figure 11-7 Comparison of Moderate Period LR and Long Period TP systems for 

WSM 80%; displacement trace and displacement histories (x and y) in 

each bearing, labeled by position (E, S, N, W) ...........................................214

Figure 11-8 Comparison of Moderate Period LR and Long Period TP systems for 

Rinaldi 88% (XY); displacement trace and displacement histories 

(x and y) in each bearing, labeled by position (E, S, N, W) ............................. 215

Figure 11-9 Comparison of the base shear coefficient for Moderate Period LR and 

Long Period TP systems in the x- and y-directions for the East bearing; 

WSM 80% and Rinaldi 88% (XY) ...................................................................216

Figure 11-10 Rotation angle at the base for Moderate Period LR and Long Period TP 

isolated configurations for (a) WSM 80% and (b) Rinaldi 88% (XY) ...............216
xx

Figure 11-11 Qualitative illustration of expected weight distribution (calculated by 

natural distribution of the weight) and observed weight distribution 

(measured in load cells) on TP isolation system .............................................217

Figure 11-12 Maximum residual displacement in the Long Period and Moderate 

Period isolation systems after each simulation ...............................................218

Figure 11-13 Displacement of the SW bearing in the Moderate Period isolation system 

over the course of a single test day, illustrating the residual displacement 

after each record ............................................................................................219

Figure 11-14 Displacement of the SE bearing in the Long Period isolation system over 

the course of a single test day, illustrating the residual displacement after 

each record ....................................................................................................219
xxi

LIST OF TABLES

Table 2-1 Yield and ultimate strength of steel from coupon tests............................................14

Table 2-2 Estimated weight in kN of the specimen by component and floor level...................16

Table 2-3 Experimentally determined natural periods and damping ratios of the fixedbase building..........................................................................................................18

Table 3-1 Required Diameter of the Lead Plug. .....................................................................28

Table 3-2 Ratio of Lead Plug Diameter to Bearing Diameter (DLP/D)......................................28

Table 3-3 Lead Rubber Bearing Properties. ...........................................................................32

Table 3-4 Cross Linear Bearing Properties.............................................................................32

Table 3-5 Predicted x, y and Vector Peak Displacement in the different Isolator for the 

Building Model Subjected to 150% Vogtle #13 Record, with Rotated 

Horizontal Components of Input Motion.................................................................. 37

Table 3-6 Common Earthquake Records Considered for Three Test Configurations. ............39

Table 4-1 Properties of load cells...........................................................................................43

Table 4-2 Vertical load on each bearing after installation. ......................................................53

Table 4-3 Simulation schedule for the TP isolation configuration............................................55

Table 4-4 Simulation schedule for the hybrid LR isolation configuration.................................56

Table 4-5 Simulation schedule for the fixed-base building......................................................57

Table 4-6 Peak acceleration of target and realized motions for the TP isolation 

configuration...........................................................................................................59

Table 4-7 Peak acceleration of target and realized motions for hybrid LR isolation and 

fixed-base configurations........................................................................................60

Table 5-1 List of ground motion (GM) simulations by number, name, abbreviation and 

the directions considered........................................................................................73

Table 5-2 Peak base shear for all simulations: total, x and y directions.................................. 76

Table 5-3 Peak base shear coefficient for all simulations: total, x and y directions. ................76

Table 7-1 Compression Shear Test Schedule......................................................................127

Table 7-2 Recorded and fitted parameters for Test A...........................................................128

Table 7-3 Characterized isolator parameters for all tests in the sequence............................128

Table 7-4 Characterized bearing parameters for each earthquake simulation. .....................132

Table 8-1 Weight and eccentricity (distance from geometric center to center of mass) 

of numerical simulation model. .............................................................................147

Table 8-2 Experimentally determined natural periods and damping ratios of the fixedbase building........................................................................................................148

Table 11-1 Ground Motions, Scale Factors and Peak Accelerations Common to 3 

Configurations......................................................................................................204

xxiii

EXECUTIVE SUMMARY

Seismic isolation systems have been proven to provide superior performance and meet

continued functionality performance objectives for many facilities around the world, and are 

thus being considered for the future generation of nuclear power plants in the United States.

Experimental simulation of a hybrid lead-rubber isolation system for a 5-story steel moment

frame was performed at Hyogo Earthquake Engineering Research Center (E-Defense) of the

National Institute for Earth Science and Disaster Prevention in Japan. The isolation system was 

developed for the seismicity of a potential nuclear site in Central and Eastern United States

(CEUS). The isolation system was tested to displacements representing beyond design basis

ground motions at the CEUS site and design basis ground motions for a Western United

States. Forces in the lead- rubber (LR) bearings were measured by an assembly of load cells.

The design of the isolation system was constrained by the experimental setup. The light axial 

loads on the system, which are not representative of a nuclear facility, necessitated the use of

a hybrid system of bearings and flat sliders, known as cross linear (CL) bearings. The CL 

bearings provide support beneath some of the columns without contributing to the system base

shear, so that the desired isolation period could be provided at the target displacement.

Additionally, the CL bearings provided substantial resistance against the tensile demands

generated by overturning as a result of the light axial loads. Thus, the suitability of a hybrid

isolation system for nuclear power plant was evaluated as part of the test program.

The following behaviors, many of which have been observed before, were observed in the 

response of LR bearings during this test program.

1. Pinching near the center of the measured bearing hysteresis loop, attributed to the small

size of the lead plug;

2. Loss of characteristic strength over the duration of an excitation, associated with heating of

the lead plug;

3. Slight fluctuation of shear force during high frequency axial force variation; thought to be

insignificant;

4. Small (negligible) permanent displacements at the end of the records;

5. Significant base rotation demands due to the inability to configure the system

appropriately for torsion;

6. No loss of shear resistance at large displacements due to the stabilizing influence of the

CL bearings;

7. Transfer of axial forces from LR bearings to CL bearings at large displacements, causing

the LR bearings to sustain tension that was not induced by overturning.

While the hybrid system overcame stability issues, the vertical force demands in individual 

bearings were difficult to predict due to lack of compliance and subsequent load transfer 

between the two types of devices (LR bearings and CL bearings). Thus, the compliance 

issues suggest that the hybrid system should not be used in nuclear facilities. If it had been

possible to design an isolation system for this experiment using only LR bearings, the response

of the bearings would have been similar but without the load transfer effects.
xxiv

A numerical simulation model was developed for the isolation system and the structure. The 

numerical simulation and its comparison to experimental data cannot be directly applied to 

nuclear facilities due to difficulties in modeling the response of the tested LR bearings, and 

differences between the tested superstructure and a nuclear facility, but still provides some

useful insight. The lead-rubber bearings were modeled with a bilinear force-displacement

relation. Due to the amplitude dependence of the bearing response, the parameters of the

bilinear model were calibrated independently for each simulation. Using the calibrated model,

the predicted displacement demand of the isolators was within 10% of the observed 

experimental displacement. Improved models that can capture the effects of heating of the lead 

core, pressure variation, and rate dependence were not used here. 

For typical XY (horizontal only) input excitation, the horizontal accelerations at the roof level 

were reduced by a factor of 10 relative to the fixed-base building. Under 3D input excitation 

which included vertical shaking, the vertical accelerations of all floor slabs were amplified by a 

factor of 4-6 relative to the input vertical excitation, but the vertical amplification factor was 

essentially the same for the isolated and the fixed-base building; thus the isolation system did

not increase vertical acceleration relative to the fixed-base building. Horizontal-vertical 

coupling was detected in the vibration modes of the structure (both fixed-base and baseisolated), which caused horizontal accelerations to be amplified somewhat in a strong 3D 

excitation compared to an XY excitation. The coupling was attributed partially to a mass

irregularity; a large supplemental mass was placed at the roof level in an asymmetric 

configuration in plan. The spectral peaks associated with the coupling modes computed from 

analysis decreased significantly when the supplemental mass in the numerical model was 

moved from the roof and distributed in a regular configuration at the base level. Other sources

of coupling in the hybrid LR isolation system were hypothesized but not confirmed. The

majority of the coupling effects were replicated by a well-crafted numerical simulation model

that accounted for slab-frame interaction and refined distribution of mass over the floor

system. The design of the base isolation system and structure should consider and

accommodate these predictable horizontal- vertical coupling effects.

All data from the experiments is permanently archived and publicly accessible in the NEES 

Project Warehouse (Ryan et al. 2013a, 2013b, 2013c).
xxv 

ACKNOWLEDGMENT

This report was prepared as an account of work sponsored by an agency of the U.S.

Government. Neither the U.S. Government nor any agency thereof, nor any of their

employees, makes any warranty, expressed or implied, or assumes any legal liability or 

responsibility for any third party’s use, or the results of such use, of any information,

apparatus, product, or process disclosed in this report, or represents that its use by such third

party would not infringe privately owned rights. The views expressed in this report are not

necessarily those of the U.S. Nuclear Regulatory Commission.

This report was based on work under the primary sponsorship of the U.S. Nuclear 

Regulatory Commission through Contract NRC-HQ-11-C-04-0067. The work was part of a 

larger collaboration between the National Science Foundation (NSF) Network for Earthquake 

Engineering Simulation (NEES) through NSF Grants CMMI-1113275 and CMMI-0721399 and 

Japan’s National Research Institute for Earth Science and Disaster Prevention (NIED). The 

NEES/NIED collaboration created the opportunity to test an additional isolation system focused 

on nuclear applications. Financial support, product donations and in-kind technical support

were provided by Earthquake Protection Systems, Dynamic Isolation Systems, Aseismic

Devices Company, Takenaka, USG Building Systems, Victaulic, Tolco, Hilti, CEMCO steel, and

Allan Automatic Sprinkler. The authors are grateful for all that led to the success of the project. 

The authors also acknowledge and thank the following individuals for reviewing and providing 

input on a draft version of the report: Andrew Whittaker, Ron Mayes, Robert Kennedy, 

Amarnath Kasalanati, and Hernando Candra.

xxvii 

NOMENCLATURES

Ab Crass-sectional area of the bearing 

acX Horizontal acceleration of the compound including top connecting plate and 

bottom half of the isolator

acZ Vertical acceleration of the compound including top connecting plate and bottom 

half of the isolator

axC Acceleration at the geometric center of the floor in x-direction

axSE X component of the horizontal acceleration at the South-East corner

ayC Acceleration at the geometric center of the floor in y-direction

aySE Y component of the horizontal acceleration at the South-East corner

b Plate width

bcf Flange width of column

c Damping coefficient of additional damper

cnn Constant dependent on nth mode shape

cs Side cover rubber thickness

d Given horizontal displacement of the bearing 

D Overall diameter

DD Design displacement

DLP Diameter of the lead plug

Dmax Maximum isolator displacement

Dmin Minimum isolator displacement

Dp Lead diameter

Dpeak Absolute peak displacement for selected cycle

DTM Maximum displacement

DXA' Displacement component at the center bearing in X direction

dxC Story drift in the x-direction at the geometric center

dxSE Story drift in the x-direction at the South-East corner

Dy Yield displacement

Dy,rub Yield displacement of rubber

DYA' Displacement component at the center bearing in Y direction

E Elastic modulus

EDC Energy dissipated per cycle

EDCbilin Energy dissipated per cycle for a bilinear force-displacement loop

EDCtest Energy dissipated per cycle for experimental data
xxviii 

eu Rubber's elongation-at-break

f Frequency

Fmax Maximum force

Fmin Minimum force

FOT Overturning force

Fpeak Absolute peak force for selected cycle

fpn Horizontal displacement at base in the nth mode shape 

fqn Horizontal displacement at roof in the nth mode shape 

Fy Yield force

G Shear modulus

g Gravitational acceleration

h Plate height

I Moment of inertia

k Horizontal stiffness

K1 Initial stiffness of bearing

K1_char Characterized initial stiffness 

K1_lead Initial stiffness of lead

K1_rub Initial stiffness of rubber

K1s Initial stiffness of the CL bearing

Kb Horizontal stiffness of low-damping natural rubber

Kd Post-yield stiffness of bearing

Kd,char Characterized post-yield stiffness

Keff Effective or secant stiffness

KH Kinematic hardening modulus

KLC Vertical stiffness of load cell

KL1 First stiffness of the lead-plug for trilinear model (Figure 7-14) 

KL2 Second stiffness of the lead-plug for trilinear model (Figure 7-14)

KL3 Third stiffness of the lead-plug for trilinear model (Figure 7-14)

Kt Tension vertical stiffness for LR bearing

Kplate Plate bending stiffness

Ktotal Total vertical stiffness

ktruss Equivalent elastic stiffness of axial element

Kv Compressive vertical stiffness for LR bearing 
xxix

Kvc Compressive vertical stiffness for CL bearing 

Kvt Tension vertical stiffness for CL bearing

Lclear Clear length between adjacent load cells

L Length of element

LA Distance between a and A' measured by the displacement transducer

m Coefficient of friction

m*

nn Modal mass of the nth mode

mc Mass of the compound including top connecting plate and bottom half of the 

isolator

MyF Yield strength of spring representing bending of the flanges

Myp Yield strength of spring representing shear behavior of panel zone

N Number of rubber layers

NLR Number of lead-rubber bearings

Pallow Allowable axial load at a lateral displacement of zero

Qd Characteristic yield strength of bearing or system

Qd,char Characterized characteristic strength

Qd,lead Characteristic strength of lead

Qd,rub Characteristic strength of rubber

QLR Total characteristic strength of lead rubber bearings

QSlider Total characteristic strength of flat sliders

R Radius of curvature of friction pendulum bearing

RcX Dynamic reaction component in x-direction at the load cell level

RcZ Dynamic reaction component in z-direction at the load cell level

RX Dynamic reaction component in x-direction at the isolator level

RZ Dynamic reaction component in z-direction at the isolator level

RZ,init Initial static vertical reaction

SF Initial stiffness of spring representing bending of the flanges

Sp Initial stiffness of spring representing shear behavior of panel zone

T Period

tbp Bottom mounting plate thickness

tcf Flange thickness of column

Td Post-yield period of bearing or system

Teff Effective period of bearing or system

tip Internal plate thickness
xxx 

Tr Total thickness of rubber

tr Layer thickness

ts Shim thickness

ttp Top mounting plate thickness

v Equivalent uniform load on the beam element

Vb Total base shear in the isolator system

Vi Shear force at the ith node

Vj Shear force at the jth node

Vp Volume of panel zone web 

W Estimated weight on each CL bearing

Wi Tributary weight supported by the ith isolator

WLRBs Total weight carried by the LR bearings

Wtotal Total weight of the structure

x Length of the gusset plate

Xa X- coordinate of node a

XA' X-coordinate of displaced node A'

y Width of the gusset plate

Ya Y-coordinate of node a

YA' Y-coordinate of displaced node A'

β Equivalent damping ratio

γ Calibration factor

Δx Displacement of center bearing in x-direction computed by data processing

Δy Displacement of center bearing in y-direction computed by data processing

ζn Additional damping ratio

μ Coefficient of friction of a sliding bearing

σu Ultimate strength

σy Yield strength

σy,LP Dynamic yield strength of the lead plug

φ Rotation of center bearing computed by data processing

φpn Horizontal displacement at base in the nth mode shape

φqn Horizontal displacement at roof in the nth mode shape

ωn Angular frequency of the nth mode

ωz,rigid Vertical frequency of a "rigid body" 
1 

1. INTRODUCTION

1.1 Potential for Seismic Isolation of Safety Related Nuclear Facilities

Seismic base isolation is a technology used to protect structures from the damaging effects of 

earthquake ground motion by decoupling the structure from the foundation through the 

incorporation of a flexible interface between the two. Flexible isolation devices are incorporated 

to shift or elongate the natural period of the structure in the horizontal direction away from the 

typical dominant energy of the earthquake, thereby significantly reducing the accelerations, 

inertial forces, and subsequently base shear demands on the structure. Increased 

displacements are accommodated in the isolation system, while structural deformations are 

substantially reduced such that the structural system practically moves “rigidly” above the 

isolators. 

Seismic isolation has been shown to successfully protect the structural system and content of 

numerous structures in prior earthquakes. Most recently, many isolated buildings were shaken 

in the 2011 Tohoku Earthquake and Tsunami, with isolator displacement demands ranging from 

20 to 40 cm (Nishiyama et al. 2011). Seismic isolated structures around the world now number 

in the several thousands, and seismic isolation systems have been incorporated into the design 

of nuclear power plants in France and South Africa (Malushte and Whittaker 2005). As a result, 

seismic isolation is being considered for the future generation of nuclear power plants in the 

United States.

The structures, systems, and components of nuclear power plants are required to be designed 

for natural phenomena (such as earthquakes) without loss of capability to perform their safety 

functions (10 CFR 50, 2011). Seismic isolation is most effective for stiff, short period structures. 

Safety related nuclear structures fit this description since their horizontal period is on the order 

of 0.1 to 0.25 sec. If designed to remain elastic under strong ground motions, a short period 

structure attracts high acceleration demands that must be accommodated by internal 

equipment. If, on the other hand, a short period structure is designed to yield, it can experience 

large ductility demands relative to a longer period structure with a comparable strength 

reduction factor (e.g. Chopra 2012). Fortunately, the flexibility of the isolation system shifts the 

natural period of the building to significantly reduce both force demands to the structural system 

and acceleration demands to internal equipment. Thus, seismic isolation could be considered 

for safety related nuclear facilities if the overall system is analyzable and the response is 

predictable.

Two major classes of isolation devices have been used in the United States: elastomeric 

bearings and friction pendulum bearings. Elastomeric bearings are composed of alternating 

layers of natural, or synthetic, rubber bonded to intermediate steel shim plates. A rubber cover 

is provided to protect the internal rubber layers and steel plates from environmental degradation 

and corrosion. Elastomeric bearings can be categorized as: (1) low-damping natural or synthetic 

rubber and (2) high-damping rubber. Low-damping natural rubber material exhibits nearly linear 

shear stress-strain behavior up to approximately 150% shear strain, wherein the horizontal 

stiffness Kb is calculated as:

b

b

r

GA K

T = (1.1) 

where G is the shear modulus, Ab is the cross-sectional area of the bearing, and Tr is the total 

thickness of rubber. A typical range of G for seismic applications is 0.4-0.9 MPa (60-130 psi). 
2 

The equivalent damping ratio, β, for low-damping natural rubber ranges between 2% and 3% at 

100% shear strain. To limit displacements across the isolation interface, external supplemental 

damping devices or alternative isolation devices are typically used in parallel with low-damping 

natural rubber bearings. A higher level of damping can be achieved through the addition of 

fillers to the rubber; however, recently such devices are rarely used in the United States. 

Another type of elastomeric bearing is the lead rubber bearing. From a construction perspective, 

lead-rubber bearings differ from low-damping natural rubber bearings only by the addition of a 

lead-plug that is press-fit into a central hole in the bearing. The lead-plug deforms plastically 

under shear deformation, enhancing the energy dissipation capabilities compared to the lowdamping natural rubber bearing. The horizontal force-deformation relationship of a lead-rubber 

bearing is characterized using bilinear behavior with an effective, or secant stiffness (Keff) 

calculated as:

(1.2) 

where Qd is the bearing characteristic strength, which is controlled by the yield strength of the 

lead in shear and the area of the lead-plug; Kd is the post-yield stiffness and d is a given 

horizontal displacement of the bearing. The vertical stiffness of all types of elastomeric bearings 

(low damping rubber, high damping rubber, and lead rubber) is typically thousands of times 

larger than the horizontal stiffness so that isolation systems composed of elastomeric bearings 

provide isolation only from the horizontal components of ground shaking.

The Friction PendulumTM (FP) bearing, developed by Earthquake Protection Systems, Inc. is a 

sliding bearing that supports the weight of the structure on a bearing that rests on a sliding 

interface. A single FP bearing consists of a single sliding interface, while a Triple PendulumTM

(TP) bearing consists of multiple sliding interfaces. The sliding interface is designed with a low 

coefficient of friction, which limits the resistance to horizontal forces. The single FP bearing 

consists of a base-plate, an articulated slider and a spherical concave dish. Under horizontal 

motion, the spherical concave dish displaces horizontally relative to the articulated slider and 

base-plate. Friction between the two surfaces provides frictional resistance and energy 

dissipation, whereas the radius of curvature of the spherical concave dish provides a restoring 

force. The shear force-horizontal deformation behavior of FP bearings is characterized using a 

bilinear relationship. The horizontal strength, or zero-displacement force-intercept, Qd, is 

controlled by the bearing material and the weight W carried by the isolators, according to: 

(1.3) 

where μ is the sliding coefficient of friction of the bearing interface. The sliding coefficient of 

friction can range between 0.03 and 0.2. The post-yield stiffness of the FP bearing is controlled 

by weight acting on the isolator and the radius of curvature, R, of the spherical concave dish. 

The TP bearing consists of four spherical sliding surfaces and three independent pendulum 

mechanisms. The internal pendulum mechanism with two concave plates and a rigid slider 

determines the response during low intensity shaking. The outer stainless steel concave 

surfaces, when designed with different curvatures and friction coefficients, provide two 

independent pendulum mechanisms that determine the response during medium to high 

intensity shaking. Both the single FP and the TP bearings provide no resistance to tensile forces 

and thus are free to uplift. In certain situations uplift in the bearings could occur, e.g. bearings 
3 

on the perimeter of slender structures or those located under braced frames. For these 

situations resistance to uplift is considered desirable.

1.2 State of Knowledge and Motivation for Full Scale Testing

The following briefly summarizes the research and development that has led to the modern 

seismic isolation systems used today. Early proof of concept earthquake simulator or “shake 

table” tests are discussed, as well as later tests that examined system level technical concerns. 

Then, analytical modeling capabilities for elastomeric isolation bearings that have been 

developed mostly in conjunction with device level testing are reviewed. Finally, the need for full 

scale earthquake simulator testing of isolated structures is addressed.

 

Initial development and proof-of-concept earthquake simulator testing was carried out in the 

United States on elastomeric and friction pendulum isolators in the 1980’s and 1990’s (e.g., 

Kelly et al. 1980a, 1980b; Kelly and Hodder 1981; Zayas et al. 1987; Mokha et al. 1988, 1990; 

Kelly and Chalhoub 1990; Constantinou et al. 1990; Al-Hussaini et al. 1994). In Japan, research 

and development was also carried out at construction companies that built several 

demonstration buildings to be tested by earthquakes (Kelly 1988). 

Earthquake simulator testing eventually progressed beyond basic validation to examine 

performance of the overall isolation system under challenging loading conditions, and 

elastomeric bearings have been tested to various limit states under dynamic loading. For 

example, a series of uplift experiments were performed on slender structures isolated with 

elastomeric bearings (Griffith et al. 1988a, 1988b), and researchers have developed and tested 

several uplift restraint devices suitable for elastomeric isolation systems (e.g. Griffith et al. 1990; 

Kasalanati and Constantinou 2005). Uplift restraint in elastomeric bearings is often provided 

through limited engagement of the elastomers in tension by providing a fully bolted connection 

detail for the elastomeric bearing. One project designed a series of experiments to drive an

isolated structure out to its ultimate capacity, including large isolator displacement demands and 

structural yielding (Clark et al. 1997). The experiment showed that design strategies can be 

adopted to ensure that the isolation system is not the weak link, and that isolators can withstand 

significant tension due to structure overturning. At least two studies performed earthquake 

simulator tests where the intensity of the excitation was increased until the bearings ruptured 

(Sato et al. 2002; Takaoka et al. 2011). A more detailed review of the history of earthquake 

simulator or “shake table” testing of isolated building structures was provided in Warn and Ryan 

(2012).

Recently, much progress has been made in understanding and modeling the macro-mechanical 

behavior of natural rubber and lead-rubber bearings. Bidirectionally coupled bilinear or BoucWen models are frequently used by commercial software to represent the hysteretic behavior of 

lead-rubber bearings. These models are incapable of representing the amplitude dependence of 

the hysteretic properties and the lateral-axial interaction of the response, which may or may not 

be significant. In lead-rubber bearings, the characteristic strength of the lead plug has been 

observed to degrade with repeated cycling (Constantinou et al. 2007), which is a result of 

viscous heating of the lead plug. Theoretical models to account for heating effects have been 

developed and experimentally verified (Kalpakidis and Constantinou 2009a, 2009b), as well as 

a numerical modeling approach for implementation in finite element software for dynamic 

analysis (Kalpakidis et al. 2010). A theory of scaling based on similitude and recommendations 

for testing reduced scaled bearings to properly account for the effects of heating was then 

developed (Kalpakidis and Constantinou 2010). Constantinou et al. (2007) also evaluated the 

effects of history of loading, aging, and environmental conditions (such as extreme temperature 

variation) on the behavior of elastomeric isolation hardware, but concluded that rational 

procedures to account for these effects in design have yet to be developed. They also 
4 

concluded that current property modification factors – used to establish probable maximum and 

minimum values of key mechanical properties over the lifetime of the bearings – are probably 

conservative.

The influence of axial load on the horizontal force-deformation behavior of elastomeric bearings

leads to complexity that can affect the analyzability of the overall system; however much recent 

research has been completed to evaluate the critical load capacity and post-buckling behavior 

of the bearings. The overlapping area approximation was developed to estimate the reduction in 

critical load capacity with increasing displacement (Buckle and Liu 1994). Several studies have 

experimentally evaluated stability and post-buckling behavior of elastomeric bearings (Buckle et 

al. 2002; Warn and Whittaker 2006; Weisman and Warn 2012; Sanchez et al. 2012). They all 

concluded that the overlapping area approximation is conservative, and improvements have 

been proposed. Experimental studies also evaluated the reduction in vertical stiffness at large 

horizontal displacements (Warn and Whittaker 2007).

A simple two-spring model that include the influence of vertical load on the horizontal response 

was developed by Koh and Kelly (1987) for linear material behavior and extended by Ryan et al. 

(2005) for nonlinear material behavior. Another extension of the two-spring model considered 

large displacement/rotation theory to predict the stable post-critical behavior that has been 

observed in test data (Nagarajaiah and Ferrell 1999). More advanced distributed spring models 

accounting for these second-order effects have been developed that are also capable of 

exhibiting zero or negative tangential horizontal stiffness (Yamamoto et al. 2009, Kikuchi et al. 

2010). 

Most of the aforementioned studies (especially earthquake simulator tests) involved reduced 

scale structural models and reduced scale isolation bearings. Despite the wealth of data on 

reduced-scale systems, the following knowledge gaps still need to be addressed.

• Dynamic test data of full-scale elastomeric bearings is not available for a system 

earthquake simulator test. A full scale system test is necessary to verify that the complex 

phenomena observed in reduced-scale bearings under realistic earthquake loading (e.g. 

velocity effects, temperature effects, horizontal and vertical interaction) are scalable to 

much larger devices. Earthquake simulator tests of a full scale building isolated with 

high damping rubber bearings and natural rubber bearings with dampers were 

performed (Sato et al. 2011), but bearing force data was not recorded as part of that 

particular program.

• Larger scale tests are needed to evaluate the influence and analyzability of complex 

isolator-structure interaction effects that are affiliated with full-scale structures. Most 

reduced scale tests are bare frame tests with supplementary mass pieces bolted to the 

frame. Such tests have not examined the influence of a composite floor system, mass

and stiffness irregularities, and nonstructural components – all of which may interact with 

the structure – on the effectiveness of the seismic isolation system.

• Similar to the previous item, earthquake simulator testing of isolated building models 

under combined horizontal and vertical loading is somewhat limited. Full scale testing 

should be performed that emphasizes realistic three-dimensional input and strong 

vertical input recorded in near-fault ground motions.

1.3 Opportunity for Testing at E-Defense 

The Hyogo Earthquake Engineering Research Center of Miki, Japan, also known as E-Defense, 

contains the world’s largest single earthquake simulator facility, and the only facility in the world 
5 

capable of reproducing three-dimensional earthquake excitation to test full scale structures

(http://www.bosai.go.jp/hyogo/ehyogo/). The 15 x 20 meter (49 x 66.6 ft) table platform is driven 

by 5 actuators in each horizontal direction and 14 actuators in the vertical direction (Figure 1-1). 

At the payload capacity of 1200 metric tons (2646 kips), the table can replicate motions with 

horizontal accelerations up to 0.9g and vertical accelerations up to 1.5g (Figure 1-1). The EDefense earthquake simulator facility can uniquely meet the testing needs for full scale isolated 

structures.

 

Figure 1-1 E-Defense earthquake simulator parameters (Source: 

http://www.bosai.go.jp/hyogo/ehyogo/) 

Collaborative research between United States and Japan investigators utilizing the E-Defense 

earthquake simulator was made possible through a Memorandum of Understanding (MOU) 

between the National Science Foundation (NSF) funded Network for Earthquake Engineering 

Simulation (NEES) program and Japan’s National Research Institute for Earth Science and 

Disaster Prevention (NIED). The NEES Tools for Isolation and Protective Systems (TIPS) 

project was funded by NSF to conduct basic and applied research to overcome some of the 

technical, cost and procedural barriers to the wider implementation of seismic isolation in the 

United States. One of the perceived barriers to implementation was the lack of full scale 

validation of seismic isolation of a realistic full scale structure to demonstrate both structural and 

content protection in very large earthquakes. Thus, the NEES TIPS project included 

collaboration with Japan to test a full scale isolated building at E-Defense.

Planning of the NEES TIPS/E-Defense test program began in earnest in May of 2010, with a 

projected test date of fall 2011. The development of the test program initially prioritized the 

testing of TP bearings for several reasons. The TP bearing has been regarded as an innovative 

isolation device that can be designed under the Performance Based Earthquake Engineering 

(PBEE) paradigm to target various levels of protection under multiple earthquakes scenarios 

(Fenz and Constantinou, 2008; Morgan and Mahin, 2011). The isolation period of TP bearings 

are insensitive to the weight of the supported structure above, and the displacement capacity of 

the device is not limited by stability considerations. This allowed the following design strategy to 

be employed: the period of the isolation system was lengthened beyond the pulse period of 

several well known and challenging near fault motions, while sufficient displacement capacity 

was provided to meet the displacement demands of the strongest long period motions that could 
6 

be found in the Pacific Earthquake Engineering Research (PEER) Next Generation Attenuation 

(NGA) database (Chiou et al. 2008). Finally, strong in-kind support from the manufacturer 

helped to overcome financial-based limitations on scope. 

The developed test program made use of an existing 5-story steel moment frame building 

structure that had been tested at E-Defense in early 2009 as part of a program on value-added 

buildings (Kasai et al. 2010). This building was stored in the fabrication yard at E-Defense from 

2009 to 2011. Over the course of 2010 and early 2011, the members of the NSF funded NEES

Research Nonstructural Grand Challenge project (http://www.nees-nonstructural.org) were 

recruited as collaborators. The test plans evolved to include installation and testing of an 

integrated interior partition wall – suspended ceiling – and sprinkler piping system on the 4th and 

5th stories of the building; and the moment frame building would be tested in the fixed-base 

configuration as well as the isolated configuration. As a nontrivial part of the program, a support 

assembly was designed for the isolated building utilizing 73 of 76 of the triaxial load cells owned 

by E-Defense to measure the shear and axial forces in each of 9 TP bearings.

The TP isolation system designed for testing at E-Defense was not ideal for a safety related 

nuclear structure, because (a) 1 meter displacement demand across the isolation interface is

rather large for a nuclear structure, (b) the design strategy using a very long isolation period 

may lead to undesirable residual displacements in the isolators, and (c) the low friction force of 

the inner sliders may activate the isolation system more than desired in small motions (say 1/6 

of design basis). The design was developed without considering the expected seismicity of a 

representative nuclear site and the needs of a safety related nuclear structure. Furthermore, 

given the history of investigation of elastomeric bearings for nuclear applications, verification of 

the stability and load-carrying capacity of elastomeric bearings at displacements representative 

of extended design basis ground motions was a top objective.

The planned test program at E-Defense offered a payload opportunity to design and test an 

additional isolation system tailored to the needs of the nuclear industry at low incremental cost. 

Thus, on June 1, 2011 (about 2.5 months before the planned start of testing at E-Defense), the 

project team, with product and in-kind support from Dynamic Isolation Systems, Inc. made the 

decision to proceed with the design and manufacture of an additional isolation system using 

lead rubber (LR) bearings to be tested at E-Defense on the potential of a future contract from 

the United States Nuclear Regulatory Commission. 

The main objective for the payload project was to evaluate the performance of the elastomeric 

isolation system designed for a safety related nuclear structure in beyond design basis

earthquake (DBE) shaking. Ground motions were developed for a representative central and 

eastern United States soil site. The test program was developed considering the performance 

objectives of ASCE 43-05 (ASCE 2008), that the isolation system has less than 1% probability 

of unacceptable performance for 100% DBE shaking and less than 10% probability of 

unacceptable performance for 150% DBE shaking. For acceptable performance, individual 

isolators are expected to (1) sustain no damage during DBE shaking; (2) sustain gravity and 

earthquake induced axial loads at 90th percentile lateral displacements consistent with 150% 

DBE shaking; and (3) have 1% or less probability of impacting surrounding structure for 100% 

DBE shaking and 10% for 150% DBE shaking. The third objective was not directly assessed in 

the test program, but relates to the predictability of the isolation system.

The E-Defense test setup provided an excellent opportunity to test an isolation system at large 

scale, but the following should be noted:

• The isolation system was to be designed for the existing (flexible) moment frame 

structure, which is unlike most nuclear structures that are extremely stiff. 
7 

• The moment-frame structure was light, which precluded the use of only elastomeric 

bearings to simultaneously provide sufficient period elongation and meet the 

displacement demands at a representative United States nuclear site. This led to the 

pairing of LR bearings with near frictionless tension-capable Cross Linear (CL) bearings

manufactured by THK, referred to as the hybrid LR isolation system. LR and CL 

bearings have been used together in over 100 projects in Japan. Thus, the hybrid 

system was found to be a reasonable solution for the test. This report examines whether 

a hybrid system is a feasible solution for a nuclear facility.

• The test setup was constrained to use load cell connection assemblies that had been 

designed for the TP isolation system. This led to the decision not to place load cells 

beneath the CL bearings because the designed connection assembly was too flexible.

• The payload project was limited to 2 days of testing with an expected limit of 7 

simulations per test day. Thus, the testing needs had to be prioritized with recognition 

that all needs could not be met.

• The objectives of other interested parties, including E-Defense collaborators, Dynamic 

Isolation Systems, Inc. and partner Aseismic Devices Co., Ltd. - provider of the CL 

bearings – were also considered in planning the test program.

• Several constraints on the test program were imposed due to safety considerations of 

the facility, such as limitations on the displacement, a slow ramp up (3-5 simulations) to 

the largest anticipated motion to ensure the target displacement limit was not exceeded, 

etc.

1.4 Report Organization

This report is organized as follows. Chapter 2 describes the 5-story steel moment frame 

building, for which the hybrid LR isolation system was designed, as well as the nonstructural 

components and building content that were monitored during the tests. The design of the hybrid 

LR isolation system, including the iterative thought process through which the final selections 

were derived, is described in Chapter 3. This chapter also explains the consideration for 

selection and scaling of ground motions during the 2 days of testing on the hybrid LR isolation

system. Chapter 4 summarizes the test program, including the assembly and connection of 

isolators and building to the earthquake simulator, instrumentation to measure structural and 

bearing responses, shaking test schedule, amplitude of realized table motions relative to the 

targets, and algorithms to generate derived data. For completeness, the shaking test schedule 

includes the test days for the TP configuration and the fixed-base configuration, although the 

test results for these systems are discussed only briefly in this report. A statistical summary of 

the test results for the hybrid LR building configuration is given in Chapter 5, where peak values 

of key response quantities are identified for every simulation. Chapter 6 examines the technical 

response and unique aspects of the hybrid LR isolation system. 

The latter part of the report furthers the understanding of the test data through development, 

validation and synthesis of a robust analytical model of the hybrid LR isolation system and

building. In Chapter 7, the analytical models for the LR bearings and CL bearings are presented, 

and the parameters of the model are calibrated to the bearing test data. Chapter 8 describes the 

analytical model of the superstructure. The analytically predicted responses of the LR bearings 

and the structural system are compared to the test data for four representative excitations in 

Chapter 9, including a synthesis of modal participation through interpretation of the floor 

spectra. Chapter 10 examines the influence of vertical excitation on the system response by 

comparing the isolator and structural responses to two pairs of ground motions that were 
8 

applied with and without vertical excitation, with further corroboration through analytical 

modeling. The response of the other building configurations (isolated with TP bearings and 

fixed-base) is briefly described and compared to the response of the hybrid LR configuration in 

Chapter 11. Chapter 12 summarizes the major findings from this test program.
9 

2. DESIGN AND CONFIGURATION OF THE 5-STORY STEEL 

MOMENT FRAME TESTBED STRUCTURE

2.1 Description of Testbed Structure

The testbed structure used in this experiment program was designed by Hyogo Earthquake 

Engineering Research Center in 2008 and used in a test in March 2009 (Kasai et al., 2010). The 

authors of this report were not involved in its design. Hereafter is description of the testbed 

structure to enable later interpretation of results. The complete structural drawings of the 

building with hybrid LR isolation system are given in Appendix A.

2.1.1 Basic Dimensions

The testbed specimen was a five-story steel moment frame building with rectangular plan

(Figure 2-1). The building was 10 x 12 m (32.8 x 39.4 ft) in plan and approximately 16 m (52.5 

ft) in height, with 2 bays in each direction. The bay widths in the long direction – 12 m or 39.4 ft

– were 7 m (23 ft) and 5 m (16.4 ft) to promote torsion, which is typical of asymmetric structures. 

Figure 2-2 shows the basic dimensions of the building and the assumed coordinate system for 

presentation of results, where the Z-axis is the vertical axis. The previous researchers chose to 

simulate a 5-story steel specimen because it is on the high side of the typical building stock in 

Japan and without added damping, would be susceptible to more significant demands than 

comparable taller buildings (Kasai et al. 2010).

Figure 2-1 The 5-story steel moment frame specimen
10

Figure 2-2 Basic dimensions of the specimen: (a) typical plan view from floors 2 to 5, 

and (b) elevation view

2.1.2 Design Spectra and Design Criteria

The design of the lateral system was based on Japanese Level II and Level III earthquake 

design spectra (Figure 2-3). Because the testbed was designed to be a “value-added” building

(Kasai et al. 2008), whose structural components and non-structural components are protected 

for major earthquakes, the story drift angle of the frame was limited to 0.005 rad for a Level II

earthquake, whereas the drift angle limit for conventional frames would be 0.01 rad. The 

structure was also required to remain elastic for a drift angle up to 0.01 rad (Kasai et al. 2008). 

2.1.3 Framing System

The lateral system was designed and detailed according to Japanese code and construction 

practice. The framing system was a three dimensional steel moment frame where the columns 

were engaged in flexure about both their principal axes. The columns were made of 35 cm x 35 

cm (13.8 in x 13.8 in) hollow box sections with thickness varying from story to story. The beams 

were either rolled or built-up I-sections. The primary beams, which were connected to the 

columns, consisted of a small-section segment at the middle and two large-section segments at

the ends (Figure 2-4). These 3 segments were all 40 cm (15.7 in) height and bolted together at 

the approximate inflection points determined from gravity loading. Connections between 

columns and beams were all fully restrained moment connections, with both flanges and web of 

the beam welded to the column. Generally, the flanges of the primary beams were haunched at 

their ends to form plastic hinges away from the columns and improve the beam-to-column 

connection. Continuity plates were also provided to protect the panel zones (Figure 2-4). 

7 m 5 m

3.85 m 3 m 3 m 3 m 3 m

1: concrete block, size 2 x 4 x 0.18 m

2: concrete block, size 2 x 2 x 0.25 m

3: concrete block, size 0.8 x 1.5 x 0.45 m

(a) (b)

7 m 5 m

5 m 5 m

1

1 2

3

Y

X
11

Figure 2-3 Design spectrum for Japanese Level II and Level III earthquakes

Figure 2-4 Beam, beam to column connection and slab

To connect the testbed to the shaking table and provide the stiffness required to “fix” the testbed 

at its base, column bases and grade girders were designed with special details. The column 

bases were detailed as steel boxes with dimension of 2.5 m x 2.5 m x 0.9 m (8.2 ft x 8.2 ft x 3 ft) 

(Figure 2-5). Vertical stiffeners were installed inside the boxes. The grade girders were bolted to 

0 1 2 3 4 5 6

0

0.5

1

1.5

2

Period T (s)

Spectral acceleration Sa (g)

Level-II

Level-III

Corrugated deck slab

Concrete protecting 

shear studs

Connection between 

segments

Large section segment

Stress haunch

Continuity plate

Small section segment
12

the column bases and were the same height as the column bases (0.9 m or 3 ft). The 

foundation framing was braced in its plane as shown in Figure 2-6. 

Figure 2-5 Column base: (a) view from top (b) view from bottom, (c) rendering of 

stiffeners

(a) (b)

(c)
13

Figure 2-6 Horizontal braces at base level

2.1.4 Slabs

The composite floor slabs were formed from 7.5 cm (3 in) high corrugated (ribbed) steel decking 

and 8 cm (3.1 in) thick normal weight concrete cast over the ribs of the deck. The corrugated 

steel deck (Figure 2-4) was 1.2 mm (0.05 in) thick and oriented parallel to the Y-direction. 

Typical reinforcement in the floor slabs was a single layer of φ1 cm @ 15 cm (φ0.4 in @ 5.9 in)

rebar in both directions placed at the slab mid-thickness. 

The roof slabs were 15 cm (5.9 in) normal weight concrete slabs cast on a 1.2 mm (0.05 in) flat 

steel deck. Reinforcement for the roof slab included layers of φ1.3 cm @ 20 cm (φ0.5 in @ 7.9 

in) rebar in each direction at the top and bottom of the slab. The roof slab was nearly twice as 

thick as the floor slabs as it was designed to carry roof mounted equipment (e.g., air 

conditioning system, water tanks) and a penthouse. 

Shear studs connected the concrete slabs to the primary beams to provide a fully composite 

connection (Figure 2-4).

2.1.5 Material Properties

The specified yield strengths of steel were 295 MPa (42.8 ksi) and 325 MPa (47.1 ksi) for the 

columns and beams, respectively. The expected ultimate tensile strengths were 400 MPa (58 

ksi) for columns and 490 MPa (71 ksi) for beams. Coupon tests showed that yield and ultimate 

strengths of the steel varied from member to member and the average ratio of measured to 

nominal strength was 1.2 (Kasai et al. 2010). Table 2-1 presents the range of observed yield

strength σy and ultimate strength σu of steel used for the beams and columns.
14

Table 2-1 Yield and ultimate strength of steel from coupon tests. 

Member σσyy(MMMMMM) σσuu(MMMMMM)

Columns 346 – 398 430 – 470 

Beams 331 – 422 510 – 557

(Source: Kasai et al. 2010)

The compressive strength of the normal weight concrete used in the slabs was 21 MPa (3 ksi); 

the measured compressive strength of standard samples was 24 MPa (3.5 ksi). The concrete 

slabs were reinforced by SD295A grade rebar. The nominal yield stress for the rebar was 295 

MPa (42.8 ksi); measured rebar strengths were unavailable. 

2.2 Non-Structural Components and Contents

Nonstructural components, including an integrated system of interior walls, suspended ceilings, 

and sprinkler piping were installed in the 4th and 5th stories, where the floor acceleration was

expected to be maximized. The overview of the nonstructural component response in both 

isolated and fixed-base building configurations is discussed briefly in this report; however, the 

detailed analysis of the nonstructural response will be reported elsewhere. Preliminary 

observations of the nonstructural response are available in Soroushian et al. (2012).

To investigate the response of non-anchored contents in the isolated and fixed-base buildings 

for different earthquake excitation, furnishings representing a hospital room on the 4th floor

(Figure 2-7) and an office room on the 5th floor (Figure 2-8) were installed in specially designed 

enclosed areas. Both rooms were 2 m x 4 m (6.6 ft x 13.1 ft) in plan and were constructed on

top of the concrete mass blocks already present on the floors (Figure 2-2(a)). Contents in the 

hospital room included a wheeled patient bed, a dresser containing medical equipment, a 

medical cart, a storage cart, IV poles, a mobile lamp, medical bottles and boxes. Many of these 

items were wheeled. The office room was furnished with desks, chairs, computer system, 

bookcases and a photocopy machine.
15

Figure 2-7 Hospital room

Figure 2-8 Office room. 

2.3 Weights

In addition to the weight of structural components, nonstructural components and contents, 

additional mass in the form of concrete and steel blocks were installed to simulate permanent 

live load. Concrete weights, whose typical size and position on the floors are shown in Figure 2-

2(a), were built as a permanent part of structure on floors 2 to 5, while steel blocks were tied to 

the roof as shown in Figure 2-9. The categorized weights of the testbed components are listed 

in Table 2-2. This information was used to design the isolation system, model the structure and 

compute inertia forces from recorded accelerations. 

In the 2009 experimental program that first used the testbed building (Kasai et al. 2010), steel 

blocks were distributed uniformly to the roof level to represent additional weight of equipment 

such as an air conditioning unit, water tank, or even a small penthouse. Each block included 

either 7 or 8 steel plates of size 2.1 m x 4.3 m x 0.025 m (6.9 ft x 14.1 ft x 0.08 ft). The roof slab 

was specifically designed to accommodate the additional weight, which summed to 820 kN (184 

kips). For this experimental program, the weight at the roof was altered from that used in 2009 

experiments to introduce additional mass eccentricity and increase torsional response; 

specifically, about one third of the steel plates were removed on the West side of the building 

(Figure 2-9). The altered weight of the steel plates for this experimental program was 535 kN

(120 kips). The intent of the added weight as designed for the original experiments was to 
16

simulate equipment weight in a typical building lacking a basement (Kasai 2011). However, this 

supplementary weight far exceeds the concentrated weight introduced by a typical single piece 

of roof mounted equipment, such as a chiller (about 80 kN or 18 kips), and thus might be 

considered atypical or even unrealistic. The issue is noteworthy because the supplementary 

weight influenced the seismic response of the building, which is discussed throughout the 

report. 

Figure 2-9 Location of steel weights at roof

Table 2-2 Estimated weight in kN of the specimen by component and floor level. 

Floor Structural Conc. 

Weight

Steel 

weight

Nonstructural Total

Roof 599 0 535 19.4 1153

5F 478 258 0 35.5 771

4F 497 268 0 16.2 781

3F 528 213 0 41.2 782

2F 527 176 0 89.6 792

Base 794 (*) 0 0 48.4 842

Sum w/ 

base 3422 914 535 250 5122

Sum w/o 

base 2628 914 535 202 4279

(*) Before the test, the weight of structural component at base was estimated at 256 kN. This low value 

did not account the weight of column bases. The total weight of the specimen corresponding to this value 

was 4585 kN.

The estimated total weight of the testbed, about 5122 kN (1151 kips), was well below the 

maximum capacity of the E-Defense earthquake simulator, which is 12000 kN (2700 kips) 

(Tagawa and Kajiwara 2007). The actual weight of the specimen measured during testing was 

5220 kN (1174 kips) as reported in Section 4.3. 

5 m

5 m 5 m

7 m

Steel 

weight

(8 plates)

Steel 

weight

(8 plates)

Steel 

weight

(7 plates)

Steel 

weight

(7 plates)
17

2.4 System Identification

The experimental response of the fixed base building to white nose excitations was analyzed 

(Sasaki et al., 2012) to find the periods and damping ratios of natural modes of the structure.

Figure 2-10 shows examples of transfer functions determined from the white noise excitation on 

the fixed-based building conducted prior to the primary earthquake excitation. The period and 

damping ratio corresponding to the fundamental response modes were evaluated by curve 

fitting theoretical transfer functions to the measured transfer functions using a least square

algorithm. Table 2-3 summarizes the periods and damping ratios of first 3 modes in both 

directions determined from this process. Rocking of the earthquake simulator has been 

observed to affect the natural frequencies and mode shapes (Kasai et al., 2011; Sasaki et al., 

2012); however, the effect of rocking was ignored in the modes presented in Table 2-3. 

With a natural period in the range of 0.65 to 0.70 seconds, the tested moment frame building is 

much more flexible than a nuclear facility, which has a natural period in the range of 0.1 to 0.25 

sec. Thus, the response of the structural system in this experiment is not representative of a 

nuclear facility. In the presentation of the superstructure response, commentary is offered to 

explain how the response of a stiffer structure might differ from the response of the test frame. 

Although the superstructure is not representative of a nuclear facility, the experiments still 

provide useful information about the dynamic response of large scale elastomeric bearings 

(designed for nuclear seismicity) in a system level test. 

Figure 2-10 Transfer function during white noise excitation in fixed-base configuration: 

(a) x-direction and (b) y-direction

(a) (b)
18

Table 2-3 Experimentally determined natural periods and damping ratios of the fixedbase building. 

White noise X White noise Y White noise 3D

Period 

(s)

Damping 

ratio (%)

Period 

(s)

Damping 

ratio (%)

Period 

(s)

Damping 

ratio (%)

Mode 1 X 0.65 3.3 n/a n/a 0.68 4.1

Mode 2 X 0.20 1.6 n/a n/a 0.21 2.0

Mode 3 X 0.11 3.3 n/a n/a 0.11 3.7

Mode 1 Y n/a n/a 0.68 2.5 0.69 3.5

Mode 2 Y n/a n/a 0.21 1.7 0.21 1.9

Mode 3 Y n/a n/a 0.11 2.6 0.11 3.6

2.5 Condition of the Testbed Prior to Testing

The testbed was built in 2008 and tested first in March 2009 equipped with several types of 

damping devices (Kasai et al. 2010). The testbed was stored outdoors and unprotected 

between April 2009 and July 2011. 

Several cracks in concrete slabs formed during the March 2009 test (Kasai et al. 2010). 

Examples of these cracks are shown in Figure 2-11. Steel beams and columns had not been 

painted and some corrosion was present in August 2011 (Figure 2-11). We did not quantify the 

effects of this damage but suspect they are insignificant.

Figure 2-11 Cracks in concrete slab and rust on steel member. 
19

3. DESIGN OF THE ISOLATION SYSTEM

3.1 Target Response of a Nuclear Site

A main objective of the test program was to demonstrate the stable response of an isolation 

system subjected to beyond design basis shaking at a representative nuclear site. Extensive 

prior work was performed by Huang et al. (2009) to characterize the site specific response 

spectra and displacement demands of representative isolation systems for three potential 

United States nuclear sites: North Anna, Vogtle and Diablo Canyon. North Anna represents a 

Central and Eastern United States (CEUS) rock site, Vogtle represents a CEUS soil site, and 

Diablo Canyon represents a Western United States (WUS) rock site. During an initial 

consultation, the external Advisory Board recommended designing an isolation system suitable 

for one of the three sites. The options were quickly narrowed down to focus on the Vogtle site 

while eliminating the other two sites from consideration. Even in beyond design basis shaking, 

the displacement demands for the North Anna site were too small to fully realize the shaking 

potential of both present isolation hardware and the E-Defense earthquake simulator facility in 

Japan. On the other hand, the displacement demands for the Diablo Canyon site were quite 

large, making it extremely difficult to size the lead-rubber bearings to provide sufficient flexibility, 

displacement capacity, and stability for the lightweight structure. Because of these difficulties, it 

is recognized that isolation of a safety related nuclear facility may be more likely to come to 

fruition on a CEUS site; thus, the Vogtle site was deemed a more appropriate selection.

The development of site specific response spectra and spectrum matched motions for the 

Vogtle site, as utilized in this study, was described in detail in Huang et al. (2009) and is hereby 

summarized. Spectra representing the design basis earthquake (DBE) were developed by 

Southern Nuclear Operating Company. The uniform risk spectrum (URS) was developed by a 

combination of probabilistic seismic hazard analysis (PSHA), site response analysis, and 

conversion of the uniform hazard spectrum (UHS) to a URS. The seismic hazard was deaggregated for a mean annual frequency of exceedance (MAFE) of 10-4 and 10-5. Spectral 

shapes were developed for both high and low frequency spectral ranges using the attenuation 

relationship of McGuire et al. (2001) for Central and Eastern United States.. Site response 

analysis was conducted to determine the amplification of rock motion to the free field surface. 

Site class factors and resulting spectra for the high and low frequency sets were merged into 

one, which led to the site specific UHS. The UHS was converted to a URS according to the 

procedure of ASCE 43-05 (ASCE 2008). These site specific spectra for horizontal and vertical 

response are shown in Figure 3-1. 

Thirty spectrum matched motions were developed using seed ground motions selected based 

on the controlling magnitude Mw and distance r pair (Mw = 7.2 and r = 130 km) for the low 

frequency spectral range. Each set of ground motions included two horizontal components and 

a vertical component. These seed motions were spectrally matched to the Vogtle site specific 

URS for the DBE to get 30 spectrum compatible motions. These motions were then amplitude 

scaled to develop maximum-minimum spectrum compatible motions. The maximum and 

minimum components consider the effects of directionality, wherein the minimum demand 

occurs at an orientation perpendicular to the maximum demand. The 5% damped response 

spectra for the 30 sets of developed maximum-minimum motions for the Vogtle site (Huang et 

al. 2009) are shown in Figure 3-2. 
20

 

Figure 3-1 5% damped URS for the DBE at the Vogtle site. (Source: Huang et al. 2009)

Figure 3-2 5% damped response spectra for the 30 sets of spectrum compatible 

maximum-minimum motions: (a) maximum component, (b) minimum 

component and (c) vertical component. (Source: Huang et al. 2009)
21

3.2 Isolation System Design

The design properties of the LR bearings to be used in the test program were selected with the 

following considerations in mind. First, as mentioned above, the isolation system was to be 

tested under beyond design basis ground motions. Thus, the general procedure utilized was to 

scale the record(s) for the Vogtle site (Huang et al. 2009 to an intensity level corresponding to 

beyond design basis, estimate the displacement demands in the isolation system, and size the 

isolators accordingly. Second, the configuration and force demands in the isolators were to be 

selected such that connection assemblies designed for the complementary tests on the TP

bearings could also be used to measure the shear and axial forces in the LR bearings. Third, 

safety requirements were imposed by the facility to protect the equipment. 

3.2.1 Estimated Displacement Demands

To obtain an approximate estimate of the isolator demands, the isolated building was modeled 

as a simple spring-mass system (rigid structure supported by a flexible isolator driven by 

bidirectional pairs of ground motions), consistent with the assumptions in Huang et al. (2009). 

The spring, or isolator, was modeled with a bi-directionally coupled bilinear force-displacement 

relationship (Figure 3-3) characterized by the post-yield stiffness Kd (corresponding to the period 

Td), and the yield strength normalized by the weight (Qd/W). The post-yield stiffness is generally 

associated with the horizontal stiffness of the rubber while the yield strength is associated with 

the shear strength of the lead plug. Additional 2% viscous damping was assigned to the isolator 

model in each horizontal direction based on the post-yield stiffness of the isolator (Huang et al. 

2009). The weight of the building was estimated to be 4540 kN (1020 kips) for design.

Figure 3-3 Idealized bilinear force-displacement relationship for LR bearings, determined 

by post-yield stiffness Kd (stiffness of rubber), characteristic strength Qd

(strength of lead plug) and initial stiffness K1 or yield displacement Dy. An 

effective or secant stiffness Keff is determined as the peak-to-peak stiffness 

based on maximum force Fmax and displacement Dmax

A subset of the parameters considered by Huang et al. (2009) was used as a starting point for 

this project, namely Td = 2 and 3 sec, and Qd/W = 0.06 and 0.09. The systems with Qd/W = 

0.06 and 0.09 are hereafter referred to as Q6 and Q9. Other values of Td, between 2 and 3 sec, 

were considered. The isolation period was desired to be greater than 2 sec to maximize the 

isolation effect and go beyond the soil-column related peak in the Vogtle spectrum just below 2 

sec. A challenge with this testbed was to provide both the flexibility and the displacement 

Lateral Displacement

Shearing force
22

capacity required at Td = 3 sec due to the relatively low weight of the building. Figure 3-4(a) 

shows the median displacement demands of the maximum-minimum motions (Figure 3-2) 

predicted for a Q6 and Q9 system at different periods, determined by statistical analysis of the 

responses of the single degree-of-freedom (SDOF) system; the values at 2 and 3 sec are 

identical to those in Huang et al. (2009). Reducing the period for the Q6 system below 3 sec did 

not reduce the displacement demand, which is likely related to the local maximum in the 

response spectrum near 2 sec. The peak displacement demand of the Q6 nonlinear system 

occurred between periods of 2.4 and 2.6 sec. For the Q9 system, the displacement demand 

decreased monotonically as the period was reduced below 3 sec. Figure 3-4(b) shows the 90% 

percentile displacement demand of the maximum-minimum motions scaled to 150% – 

representative of beyond DBE shaking – for a Q6 and Q9 system. For this case, reducing the 

period below 3 sec caused the 90th percentile displacements to increase for both Q6 and Q9 

systems. Based on these observations, a target period of 3 sec was selected for initial design 

and both Q6 and Q9 systems were considered.

Figure 3-4 Displacement demand representing (a) median response to maximumminimum motions and (b) 90% percentile response to 150% maximumminimum motions for Vogtle site

3.2.2 Configuration of Lead-Rubber Bearings

As mentioned previously, the testbed was a 2 bay frame in each direction with 9 columns. The 

configuration of columns at the base labeled by direction coordinates is shown in Figure 3-5. 

Forces in LR bearings could potentially be obtained by bolting the bearings to connection 

assemblies. The connection assemblies, which were designed for the TP isolation system,

consisted of 7 to 9 load cells sandwiched between two steel plates (e.g. Figure 3-6). Further 

details of the connection assemblies are provided in Chapter 4.
23

Figure 3-5 Drawing of testbed base plan with column labels (N = North, E = East, S = 

South, W = West, C = Center)

Figure 3-6 Illustration of a connection assembly with triaxial load cells to measure 

bearing forces

The target design parameters of the system could only be achieved with a small number of LR 

bearings. Thus, a design using four LR bearings was proposed. Given that LR bearings would 

not be placed beneath every column, two alternatives were considered. In Configuration Option 

1, LR bearings were to be placed at the four corner columns (SE, SW, NE and NW) with no 

isolators beneath the remaining 5 columns (Figure 3-7(a)). In Configuration Option 2 (Figure 3-

7(b)), LR bearings were to be placed beneath the four edge columns (S, E, W and N), and 

complementary flat sliders were to be placed beneath the remaining 5 columns, comprising a 

hybrid isolation system as introduced in Chapter 1. (Note that the selected CL bearings were 

one of several types of flat sliders that could have been used for this purpose.) The weight 

supported by each isolator based on tributary load calculation is indicated for each configuration 

option. Configuration Option 1 was preferred since the LR bearings were to carry the total 

E

S

N

C W

SE

NE

SW

NW

Shake table

Load cell

Bottom 

connecting 

plate
24

weight of the building, leading to a good test of the isolator stability, which was related to the 

test objectives. Configuration Option 1 was ultimately eliminated because the base diaphragm 

was not believed to be sufficiently stiff to suppress bouncing of the unsupported columns, and 

the expected loads on individual isolators exceeded the capacities of the designed connection 

assemblies (Figure 3-6). 

Thus, Configuration Option 2 was selected. The LR bearings were located at the edge columns 

in Configuration Option 2 (Figure 3-7(b)) where they were expected to carry larger axial forces, 

which was desirable for a better test of the stability of the bearings at large displacements. A 

drawback to this arrangement was that it decreased the torsional resistance of the isolation 

system from Configuration Option 1. We note that a typical isolation solution for a large structure 

or nuclear power plant would not be affected by these constraints.
25

Figure 3-7 (a) Configuration Option 1 with LR bearings beneath 4 corner columns, (b) 

Configuration Option 2 with LR bearings (circles) beneath 4 edge columns 

and flat sliders (squares) beneath remaining columns. The supported 

weights (in units of kN) at the isolators based on tributary loads are also 

indicated

 (b) 

SE S

E

NE N NW

W

SW

C

385 605 185

640 1020 400

380 620 305

SE SW

NE NW

1164

1035

1171

1170

 (a) 
26

3.2.3 Selection of the Bearing Dimensions

Based on the target period Td = 3 sec and a displacement demand of 60 cm (24 in) for a Q6 

system and 50 cm (20 in) for a Q9 system (Figure 3-4(b)), three alternative isolators were 

proposed by Dynamic Isolation Systems, who provided the bearings for this project. The 

parameters for each alternative are shown in Figure 3-8, where D is the overall diameter, N is 

the number of rubber layers, Pallow is the allowable axial load at a lateral displacement of zero, 

and Dmax is the maximum displacement capacity of the bearing at the anticipated axial load 

demand. All options assumed 0.6 cm (0.236 in) thick rubber layers and a shear modulus G = 

0.41 MPa (0.06 ksi). The post-yield stiffness Kd was calculated from

b

d

r

GA K

T = (3.1)

where Ab is the area of the bearing, and Tr is the total thickness of rubber. Pallow was computed 

as the maximum axial load of the bearing in the undeformed configuration divided by a factor of 

3, which is standard industry practice in the United States to provide a high margin of safety 

under gravity loads. Figure 3-9 illustrates the theoretical axial load capacity (computed as a 

minimum of buckling, elastomer limit or stress limit) for each proposed bearing design as a 

function of lateral isolator displacement (Constantinou et al. 2007). The peak axial load demand 

was estimated for the trial calculations as the peak dead weight supported by any isolator 

(based on tributary area) increased by a factor of 1/3 for overturning: 850 kN (191 kip). For the 

trial design the displacement capacity was estimated as the displacement at an axial load of 850 

kN (191 kip) on the axial force-lateral displacement curve (Figure 3-9). 

Figure 3-8 Proposed design options for LR bearings 

 LR Option 1 LR Option 2 LR Option 3

D = 58.4 cm D = 63.5 cm D = 69.9 cm 

N = 34 N = 34 N = 40

Td = 2.87 sec. Td = 2.67 sec. Td = 2.62 sec.

Dmax = 50 cm Dmax = 50 cm Dmax = 60 cm

Pallow = 2380 kN Pallow = 3327 kN Pallow = 4197 kN
27

Figure 3-9 Axial force capacity of proposed LR bearings versus lateral displacement

LR Option 1 was the most flexible of the three (Td = 2.87 sec) and it nearly met the target 

period, but its displacement capacity was limited. The target displacement could only be 

achieved for Q9. LR Options 2 and 3 provided nearly the same effective period, but LR Option 3 

had a substantially higher displacement capacity, which was achieved by increasing the 

diameter and height of the bearing. LR Option 3 was preferred to LR Option 2.

3.2.4 Selection of the Lead Plug Dimensions

Initially, lead plugs were sized for Q6, Q9 and an intermediate option Q7.5 (Qd/W = 0.075), 

where contributions to the zero-displacement force intercept were to be made by flat sliders with 

a friction coefficient μ = 0.06 and the lead plugs in the LR bearings. Recall that LR bearings 

were to be installed under the edge columns (columns S, E, W and N in Figure 3-7(b)) and flat 

sliders were to be installed under the remaining 5 columns. The sliders alone would have 

provided a yield force of approximately 120 kN (27 kips) based on the tributary weight and a 

coefficient of sliding friction of 0.06. The lead plugs were sized to provide the remainder, based 

on the following equation:

( )

LR Slider

2

LR ,LP LP 4

d

y

Q Q WQ W

N D σ π

  = −    

= 

 (3.2)

where QLR and QSlider are the characteristic strength of the LR bearings and flat sliders, 

respectively. NLR is the number of LR bearings (4), σy,LP is the dynamic yield strength of the lead 

plug and DLP is the diameter of the lead plug. For this calculation, σy,LP was taken as 7.94 MPa 

(1.15 ksi), which is the value recommended by Dynamic Isolation Systems for their products. 

Table 3-1 lists the required diameter of the lead plug for Q6, Q7.5 and Q9, tabulated from 

Equation (3.2). The required diameter of the lead plug increased by about 33% (from 7.7 cm 

(3.0 in) to 10.6 cm (4.2 in)) from a Q6 to a Q9 design.
28

Table 3-1 Required Diameter of the Lead Plug. 

Label 

Yield 

strength 

σy,LP in 

MPa (ksi)

Strength 

required Q 

in kN (kip)

Strength 

required per 

bearing 

in kN (kip)

Area lead 

plug in cm2

(in2

) 

Diameter 

lead plug 

DLP in cm 

(in)

Q6 7.94 (1.15) 146.8 (33) 36.7 (8.25) 46.2 (7.2) 7.7 (3.0)

Q7.5 7.94 (1.15) 213.5 (48) 53.4 (12.00) 67.2 (10.4) 9.3 (3.6)

Q9 7.94 (1.15) 280.2 (63) 70.0 (15.75) 88.5 (13.7) 10.6 (4.2)

If the diameter of the lead plug is small, then the hysteresis may be pinched as seen in later 

figures. Importantly, if the plug is too small, little energy dissipation is achieved. If the diameter 

of the lead plug is too great, then the isolator may not provide sufficient confinement of the core 

during repeated cycling. The rule of thumb, based on years of experimentation and analysis is 

that the ratio of the diameter of the lead plug to the bonded diameter of the bearing is between 

1/6 and 1/3. Table 3-2 presents the ratio of DLP/D for the different combinations of bearing and 

lead plug sizes. For most combinations, the ratio did not meet the minimum of 1/6. The 

diameter ratios were lowest for LR Option 3, which provided the largest displacement capacity.

Based on Table 3-2, the combination of LR Option 3 and Q9 almost met the minimum diameter 

ratio requirement, but the displacement demand would be lower for a Q9 design, such that the 

provided capacity of the bearing would not be fully utilized in the test. Thus, an alternative low 

friction slider was considered, as described in the next section. 

Table 3-2 Ratio of Lead Plug Diameter to Bearing Diameter (DLP/D). 

Bearing 

Diameter

Q6

(DLP = 7.67 cm)

Q7.5

(DLP = 9.25 cm)

Q9

(DLP = 10.62 cm)

LR Option 1 

(D = 58.4 cm) 0.13 0.16 0.18

LR Option 2 

(D = 63.5 cm) 0.12 0.15 0.17

LR Option 3 

(D = 69.9 cm) 0.11 0.13 0.15

3.2.5 Cross Linear Bearings 

The cross-linear (CL) bearing manufactured by THK allows nearly resistance-free linear motion. 

The LM Guide technology allows free rolling motion of a weight supporting part on a rail, where 

the part and the rail are internally separated by recirculating ball bearings. The CL bearing uses 

two orthogonally mounted LM Guide assemblies (Figure 3-11). The CL bearing can be 

combined with traditional isolation devices to support the weight of the building without 

increasing the total base shear of the isolation system. The coefficient of friction of the sliders 

varies from 0.48-0.62%, leading to a negligible contribution to the base shear for a reasonable 

value of the supported weight. 

Making use of the CL bearings means that QSlider in Equation (3.2) can be taken to zero; thus 

allowing the lead plugs to provide the yield strength in entirety. Opting for a Q6 design, the 

required diameter of the lead plug was 10.16 cm (4.0 in), which provided the estimated total 

yield strength of 267 kN (60 kip) or 66.75 kN (15 kip) per bearing. 
29

Besides its low friction sliding capability, the CL bearing provides significant resistance to tensile 

forces. A more careful evaluation of overturning on the hybrid isolation system suggested that 

the LR bearings alone would be unable to resist the tensile demands. The tensile force

demands in the system at a displacement of 60 cm (24 in) were estimated as follows. First, the 

total base shear Vb in the isolation system was estimated as:

Figure 3-10 Photo of installed CL bearing illustrating orthogonal LM guides assemblies 

on top and bottom

max ( )

4(66.75 kN 0.65 kN/mm 600 mm)

1827 kN (411 kip)

V N Q KD b LR d d = +

=+ ⋅

=

 (3.3)

The post-yield stiffness Kd and strength Qd assumed LR Option 3 with a lead plug diameter DLP

= 10.16 cm (4 in). The base shear was assumed to act at 9 m, which is about half the height of 

the building (Figure 3-11). The overturning moment generated by the base shear was balanced 

by an overturning force FOT times the shortest base dimension of the building (10 m). Thus, FOT

was calculated as 9/10 of the base shear Vb (FOT = 1644 kN or 370 kip). The overturning 

demand was assumed to be carried by tension on the more lightly loaded South side of the 

building (Figure 3-7(b)). The total tension T carried by the three isolators (SE, S and SW) was 

computed as: 

1644 kN (385+605+185)kN

469 kN (105 kip)

OT i i

TF W = −

= −

=

∑ (3.4)

where Wi is the tributary weight supported by the ith isolator, summed over the SE, S and SW 

isolators. Without CL bearings, the 469 kN (105 kip) of tension would be carried by a single LR 

bearing. Since CL bearings were utilized, the CL bearings were expected to carry the 
30

overturning induced tension and each CL would be subjected to about 235 kN (53 kip) tension. 

In reality, the tensile demands may not be equally balanced by the CL bearings, since the SW 

bearing carries significantly less weight than the SE bearing according to tributary area (Figure

3-7(b)). 

Equation (3.4) suggested that peak compressive force on a given side of the building could 

increase by a factor of 2 or more due to overturning. The initial estimate of axial force demand 

(850 kN or 191 kip) used to estimate the displacement capacity of the LR bearings would then 

be unconservative. However, the CL bearings were much stiffer in compression than the LR 

bearings, and the vertical movements of LR and CL bearings were coupled together by the 

rigidity of the base diaphragm. This base diaphragm constraint was expected to prevent

individual LR bearings from shortening or buckling, thus enhancing the overall stability of the 

isolation system such that the projected axial force limits would not be relevant. The interaction 

and load transfer between LR bearings and CL bearings is a unique aspect of this isolation 

system, which is evaluated extensively later in this report. The suitability of the tested hybrid 

system for nuclear facilities is also evaluated based on a synthesis of the experimental data. 

 

Figure 3-11 Total base shear (Vb) and overturning forces (FOT) acting on an elevation

view of the testbed building

3.2.6 Summary of Design Properties 

The final hybrid LR isolation system design included four LR bearings and five CL bearings. The 

configuration of the bearings (LR bearings at edge columns and CL bearings at center and 

corner columns) is shown in Figure 3-12, along with photographs of the devices taken prior to 

installation. LR Option 3 was used for the LR (D = 69.9 cm or 27.5 in, 40 rubber layers, post 

yield stiffness Kd = 6.5 kN/cm or 3.71 kip/in and Td = 2.6 sec) and the lead plug was sized for Q6 

(DLP = 10.16 cm or 4 in). The estimated displacement capacity of the LR bearings based on a 

stability limit was 60 cm (23.6 in). The CL bearings were designed with a displacement capacity 

of 60 cm (23.6 in) imposed by a low force capacity stopper at the end of travel in each 

perpendicular direction that was not intended to be reached. 

Vb

FOT FOT
31

Figure 3-12 (a) Final plan drawing of the hybrid LR isolation system, (b) photo of LR

bearing, and (c) photo of CL bearing

The dimensions and target stiffness and strength parameters of the LR bearings are listed in 

Table 3-3. All parameters were provided by the manufacturers. Several of the parameters are 

modeling parameters recommended for a bilinear representation of the force-deformation 

relation, as shown in Figure 3-3. The bearings were tested by Dynamic Isolation Systems prior 

to shipment to E-Defense. Force-deformation characterization was generated for cycles of 

amplitude 30 cm (11.8 in), 50 cm (19.7 in) and 65 cm (25.6 in) at different axial loads. Pseudostatic tests were performed and dynamic material properties were not provided. The results of 

this characterization are reported in Chapter 7.

The main properties of the CL bearings are listed in Table 3-4. The vertical stiffness of the CL 

bearings in compression and tension was provided by Aseismic Devices Co. Design drawings 

and specification sheets for both LR and CL bearings provided by the manufacturers are 

included in Appendix B.

(a) (b)

(c)
32

Table 3-3 Lead Rubber Bearing Properties. 

Bearing Dimensions 

Overall Diameter, D = 69.85 cm (27.5 in)

Number of Rubber Layers, N = 40

Lead Diameter, Dp = 10.16 cm (4.0 in)

Shim Thickness, ts = 0.3 cm (0.1196 in)

Layer Thickness, tr = 0.6 cm (0.236 in)

Side Cover Rubber Thickness, cs = 1.27 cm (0.5 in)

Top Mounting Plate Thickness, ttp = 2.54 cm (1in)

Bottom Mounting Plate Thickness, tbp = 2.54 cm (1in)

Internal Plate Thickness, tip = 2.54 cm (1in)

Isolator Properties

Effective Period, Teff = 2.27 sec

Post Yield Period, Td = 2.62 sec

Design Displacement, DD = 30 cm

Maximum Displacement, DTM = 60 cm

Post-Yield Stiffness, Kd = 6.5 kN/cm (3.7 kip/in)

Initial Stiffness, K1 = 65 kN/cm (37 kip/in)

Characteristic Strength, Qd = 65.7 kN (14.8 kip)

Yield Force, Fy = 73 kN (16.4 kip)

Yield Displacement Dy = 1.128 cm (0.44 in)

Vertical Stiffness, Kv = 15000 kN/cm (8566 kip/in)

Shear Modulus, G = 0.414 MPa (0.06 ksi)

 Rubber Ultimate Strain (at-break), εu = 5.5

Table 3-4 Cross Linear Bearing Properties. 

Isolator Properties 

Coefficient of friction, μ = 0.48%-0.62%

Yield displacement, Dy = 0.01 cm

Compressive Vertical Stiffness, Kvc = 34710 kN/cm (19821 kip/in)

 Tension Vertical Stiffness, Kvt = 2450 kN/cm (1399 kip/in)

3.3 Selection and Scaling of Ground Motions

The test program for the hybrid LR isolation system was developed based on the 

recommendations and interests of several different parties, including the research sponsor, the 

Advisory Board members, the manufacturers, the project team and Japanese collaborators. 

Two days of testing were planned for the hybrid LR isolation system configuration, where 7 

independent trials (in extreme cases 8) could be executed each day. 

Demonstrating the stable performance of elastomeric isolation systems in design or beyond 

design basis earthquakes is an important step to enable the use of base isolation for safety 

related nuclear structures. The seismic hazard at the Vogtle site is well known to the nuclear 

engineering community, and thus a record representing the seismic hazard at the Vogtle site

was prioritized for the test program. A Vogtle record was sought from among the set of 30 

maximum-minimum spectrum compatible ground motion pairs developed by Huang et al. (2009) 

that would produce a peak LR bearing displacement demand of about 60 cm (23.6 in) when 

scaled to 150%, in accordance with the beyond design basis target. Extensive pre-test analysis 
33

was conducted to identify the best Vogtle record for this purpose. While the isolation system 

was designed specifically for beyond design basis motions at the Vogtle site, its capabilities also 

permitted the system to be subjected to a Diablo Canyon record, representative of a WUS rock 

site, scaled to about 100%. The Diablo Canyon record was selected from a set of 30 maximumminimum spectrum compatible ground motion pairs developed by Huang et al. (2009) using a 

procedure similar to the one described for the Vogtle site. The sponsor requested that the 

maximum displacement demands on the LR bearings be imposed in as few trials as possible so 

that the bearings were in their virgin state. The performance of bearings made from natural 

rubber has generally been found to be stable after repeated cyclic testing, although the 

hysteresis of LR bearings can be affected by heating of the lead plug and strain hardening 

induced by cumulative travel effects (Constantinou et al. 2007). An additional objective evolved 

from these considerations, which was to repeat one of the early trials at the end of the test 

program to evaluate the consistency of the LR bearing response.

Once the sponsor objectives had been met, other objectives could be entertained. For example, 

the project team aimed to identify a service level, design level, and maximum considered 

earthquake (MCE) level motion to be replicated on all three building configurations (TP isolation 

system, hybrid LR isolation system, and fixed base). The Japanese collaborators aimed to 

identify a strong Japanese motion that could be replicated on all three building configurations. 

Dynamic Isolation Systems requested a sine wave characterization test to be repeated at the 

beginning and the end of the test program. 

The test facility imposed additional constraints on the test program based on safety 

considerations, which are described next.

3.3.1 Imposed Limitations for Safety

The following limitations related to target demands and test sequence were imposed.

1. The target displacement demand of the LR bearings was limited to 55 cm (21.6 in). In 

initial discussions, Japanese collaborators imposed a displacement limit equal to the 

design displacement of 30 cm (11.8 in), which was comparable to displacements 

permitted in previous tests at E-Defense. A compromise was reached after a) it was 

demonstrated that shear rupture was not expected before displacements of 80-90 cm 

(31.5-35.4 in), b) it was demonstrated that the CL bearings would stabilize the entire

isolation system, and c) Dynamic Isolation Systems agreed to in-house characterization

tests of the LR bearings to displacements of 65 cm (25.6 in) prior to shipment of the 

bearings to Japan. 

2. The target displacement demand of the CL bearings was limited to about 40 cm (15.7 in) 

in each of the x and y perpendicular directions, which is a factor of safety of about 1.5 

relative to the displacement limit of the CL bearings. This agreement was reached after 

Aseismic Devices Co. agreed to add a safety stop at the end of travel in each direction. 

The safety stop was not intended to stop the momentum of the building if a high impact 

collision of the sliders with the safety stop were to occur. 

3. The largest displacement was to be approached over a series of 3 or 4 incremented

trials that gradually increased the intensity of the earthquake shaking. This incremental 

approach was intended to validate the analytical model and allow adjustment to the 

intensity of the largest imposed record as necessary. This constraint conflicted with the 

objective to impose the largest intensity record early in the test sequence, but could not 

be avoided.
34

3.3.2 Vogtle and Diablo Canyon Motions

Substantial effort was expended to identify the best Vogtle record and best Diablo Canyon 

record for the testing program. Six Vogtle records were considered; these records were 

identified by evaluating the peak bidirectional displacement demand of the SDOF system to all 

30 pairs of Vogtle records scaled by 150%, and selecting those that predicted a peak 

displacement closest to 55 cm (21.6 in). Displacement traces (x vs y-direction displacement) 

and displacement histories for the 6 records that were considered are shown in Figure 3-13 and 

3-14. Preference was given to the records that included multiple cycles of large displacement 

(Figure 3-14), and followed a partially circular trace rather than a linear trace in a given direction 

(Figure 3-13). Vogtle #13 and Vogtle #9 were considered to meet these criteria better than the 

other records.

Figure 3-13 Calculated displacement trace of the isolation system for 6 Vogtle motions 

by SDOF analysis

-60

-40

-20

0

20

40

60

Vogtle #8

-X Displacement (cm)

Vogtle #9 Vogtle #13

-50 0 50

-60

-40

-20

0

20

40

60

Vogtle #17

-50 0 50

Vogtle #27

Y Displacement (cm)

-50 0 50

Vogtle #29
35

Figure 3-14 Calculated displacement histories of the isolation system in x and ydirections for 6 Vogtle motions by SDOF analysis

Next, the records were applied to the isolated building model that was developed in OpenSees 

(described in Chapter 8) to obtain a more accurate assessment of the displacement and force 

demands. In general, the peak displacement demands of the isolators in the building model 

were somewhat larger than in the SDOF model. Upon examining the data, two sources of 

discrepancy were identified. First, the building model did not contain viscous damping alongside 

the isolation system. Displacements in the SDOF model were re-evaluated after removing this

damping (2% of critical calibrated to the isolator post-yield stiffness), and they increased 

substantially, which indicated that the response was sensitive to the seemingly small additional

damping. Although some viscous energy dissipation may be present, the damping was removed 

from the SDOF model to err on the side of conservatism. 

Second, substantial rotational demands at the base level were predicted by the analysis of the 

building model, which caused amplification of the displacement on one side or corner of the 

building compared to the other. To illustrate this, the displacement traces of each bearing are 

presented in Figure 3-15. Due to the rotation, the peak displacement in one bearing was 

predicted to be 54 cm (21.3 in), while the peak displacement predicted in the opposite corner 

was only 46 cm (18.1 in). The project team was skeptical about the significant amount of torsion 

predicted by the analysis, and experimented with the modeling assumptions to develop 

confidence in the prediction and possibly identify a cause. Several alternative assumptions were 

considered, including bearing placement at the corners rather than on the edges, and 

accounting for the rotational stiffness of both the LR and CL bearings. None of the modifications 

significantly altered the amplitude of the rotational demands, and the experimental data later 

validated the torsion predicted by numerical simulation (see Chapters 5 and 9). The rotational 

demands observed in the isolation system resulted from limitations on the number and 

placement of LR bearings for the testbed structure. In a large building or safety related nuclear 

structure with hundreds of isolators, isolation system asymmetries and rotational demands could 

be eliminated or minimized by strategic placement of bearings with and without lead plugs.

-60

-40

-20

0

20

40

60

Vogtle #8

Displacement (cm)

Vogtle #9 Vogtle #13

0 5 10 15 20 25

-60

-40

-20

0

20

40

60

Vogtle #17

0 5 10 15 20 25

Vogtle #27

Time (sec)

0 5 10 15 20 25

Vogtle #29

x-dir.

y-dir
36

With the exception of Vogtle #13, the peak displacements predicted by the building model 

exceeded those predicted by the SDOF model without viscous damping. Thus, Vogtle #13 was 

selected since the displacement predicted by the building model was closest to that predicted by 

SDOF model for the desired scale factor of 150%.

Figure 3-15 Predicted displacement trace of each isolator for the OpenSees building 

model subjected to 150% Vogtle #13

While the LR bearings could be subjected to displacements of up to 55 cm (21.7 in), the CL

bearings were not permitted to exceed displacements of 40 cm (15.7 in) in the x and ydirections. Thus, we proposed to rotate the components of the input ground motion such that 

the peak displacement demand occurred at approximately 45 degrees, which would 

simultaneously minimize the demands in x and y-directions and maximize the vector 

displacement. To determine the rotation of the input motion, the building model was analyzed to 

the Vogtle #13 input excitation rotated at increments of 11.25 deg. Thus, rotated inputs at 0, 

11.25, 22.5, 37.75, and 45 degree were considered. Based on this analysis (summarized in 

Table 3-5), a rotation angle of 11.25 degrees was selected. The anticipated peak displacement 

demand in any LR bearing for Vogtle #13 rotated by 11.25 degrees and scaled to 150% was 49 

cm (19 in), while the peak displacement in any CL bearing in the x or y direction was 40 cm (16

in). Adjustments to the scale factors were made on the day of testing, and the actual peak scale 

factor applied was 175%. The complete final schedule of simulations actually conducted is 

summarized in Chapter 4.

-50

0

50

46.9 cm 43.6 cm

Vogtle #13 - 150%

45.9 cm

-50

0

50

49.8 cm

-X Displacement (cm)

43.4 cm 44.4 cm

-50 0 50

-50

0

50

54.2 cm

-50 0 50

45.3 cm

Y Displacement (cm)

-50 0 50

43.5 cm
37

Table 3-5 Predicted x, y and Vector Peak Displacement in the different Isolator for the 

Building Model Subjected to 150% Vogtle #13 Record, with Rotated Horizontal 

Components of Input Motion. 

Bearing Rotation X - Peak Y - Peak Vector Peak

Angle Displacement Displacement Displacement

(degrees) (cm) (cm) (cm)

Lead 

Rubber

0 34 45 50

11.25 39 40 49

22.5 44 33 49

37.75 47 27 50

45 50 21 51

Cross 

Linear

0 34 45 54

11.25 39 40 55

22.5 44 33 54

37.75 47 27 53

45 50 21 52

Figure 3-16 Predicted displacement trace of each isolator for the OpenSees building 

model subjected to 150% Vogtle #13, with input ground excitation

components rotated by 11.25 degrees

-50

0

50

45.8 cm 44 cm

Vogtle #13 - 150% Rotated

45.3 cm

-50

0

50

49.4 cm

-X Displacement (cm)

44.2 cm 44.1 cm

-50 0 50

-50

0

50

54.6 cm

-50 0 50

47.2 cm

Y Displacement (cm)

-50 0 50

44.1 cm
38

A similar process was used to select the input motion to represent the design ground shaking at 

the Diablo Canyon site, with a target scale factor of 100%. The Diablo Canyon #15 record was 

selected for the test program. The predicted displacement trace of the isolators for Diablo 

Canyon #15 scaled to 100% is shown in Figure 3-17. Rotation was not required for this input 

motion since the vector displacement was approximately maximized without rotation. The actual 

scale factor applied to Diablo Canyon #15 during testing was 95%.

Figure 3-17 Predicted displacement trace of each isolator for the OpenSees building 

model subjected to 100% Diablo Canyon #15

3.3.3 Remainder of the Test Program

As mentioned previously, the objectives of the sponsor were prioritized in the planning of the 

test program. These objectives were to simulate the response of LR bearings to motions 

representing the seismicity at a potential nuclear site, and impose design basis and extended 

design basis demands on the system. Once these objectives had been satisfied, objectives of 

other interested parties could be entertained. 

For comparative purposes, the project team (United States and Japan collaborators) proposed 

to include in the test program 3 ground motions, one each representing a service level, design 

level and MCE as defined by the United States building code (ASCE 2010), that would be 

commonly applied to each of the three building configurations. The assumed seismic hazard 

associated with these events is presented in Dao (2012). In addition, Japan side collaborators 

-50

0

50

61.8 cm 51.2 cm

Diablo #15 - 100%

46.4 cm

-50

0

50

53.4 cm

-X Displacement (cm)

40.7 cm 34.6 cm

-50 0 50

-50

0

50

47 cm

-50 0 50

36.7 cm

Y Displacement (cm)

-50 0 50

32.6 cm
39

requested that a large motion recorded during a Japan earthquake be commonly applied to 

each of the three building configurations. The preliminary selections are shown in Table 3-6.

During the testing of the TP isolation configuration, which was chronologically first in the 

sequence, the response of the building was particularly affected by the strong vertical excitation 

of the 1994 Northridge recorded at Rinaldi Receiving Station. This excitation was not part of the 

planned test program for the hybrid LR isolation or fixed-base configurations. However, late 

modifications to the planned test program were accommodated to repeat this excitation as a 3D 

excitation and an XY excitation (omitting the vertical component) in each building configuration 

to better comprehend the response of isolated and non-isolated buildings to strong vertical 

excitation. 

In the end, not all objectives were met due to safety considerations (see Table 3-6) and 

compromises were made. The imposed safety limits were analytically predicted isolator 

displacements ≤ 55 cm for the hybrid LR isolation system and analytically predicted structural 

drift limits ≤ 1.2% for all configurations. A suitable MCE level earthquake that met the safety 

limitations for the hybrid LR isolated and fixed-base configurations could not be identified, and

the MCE comparison was removed from the test program. The selected design event (El 

Centro) was never applied to the fixed-base configuration, and the selected Japan motion 

(Iwanuma) as well as Rinaldi were applied to the fixed-base building at reduced scale factors in 

the horizontal direction.

Table 3-6 Common Earthquake Records Considered for Three Test Configurations. 

Objective Earthquake Record Scale 

Factor

Simulation Considered Safe?

TP 

Config.

Hybrid LR

Config.

FixedBase 

Config.

Service 

Earthquake 1987 Superstition Hills, 

Westmorland Sta. (3D) 80% Yes Yes Yes

Design 

Earthquake 1940 Imperial Valley, El 

Centro Sta. (3D) 130% Yes Yes Safety questions

MCE 

Earthquake

1994 Northridge, Sylmar 

OR 1995 Kobe, Takatori

(3D)

100% Yes

No, safety 

imposed

displacement 

limit

No, safety 

imposed 

story drift 

limit

Japan 

Earthquake 2011 Tohoku Earthquake, 

Iwanuma (XY) 100% Yes Yes No, Scaled to 70%

XY vs 3D 

Input 

Comparison

1994 Northridge, Rinaldi 

Rec. Sta. (XY) 88% Yes Yes No, Scaled to 35%

XY vs 3D 

Input 

Comparison

1994 Northridge, Rinaldi 

Rec. Sta. (3D) 88% Yes Yes

No, Scaled 

to 35% 

(88% in 

vertical)

41

4. EXPERIMENTAL PROGRAM

4.1 Design of Connection Assembly

A plan was developed and executed for securely connecting the isolation devices (both LR 

bearings and CL bearings) to the structure and to the earthquake simulator. As described in 

Chapter 3, the connection design for the hybrid LR isolation system made use of connection 

assemblies, each consisting of a layer of load cells sandwiched between a square or 

rectangular base plate and a hexagonal shaped top plate, that were designed for the TP 

isolation configuration. Details of the assembly design calculations and supporting finite element 

analysis were described in Dao (2012). The assemblies were used at the N, S, E and W column 

locations to measure the forces in the LR bearings. 

The connection assemblies were not used at CL bearing locations, for several reasons. First, 

the axial force demands on the CL bearings were expected to be high since the CL bearings 

would carry all overturning induced axial forces. Second, Aseismic Design Corporation, the 

provider of the CL bearings, calculated that the supporting plates of the connection assembly 

were too flexible to prevent rotation (bending) of the bearings about the horizontal axes, and 

thus ensure their proper function. Although very small, the contribution of the CL bearings to the 

total base shear could not be measured.

Three distinctly configured load cell connection assemblies had been devised for the TP 

isolation configuration tests according to expected tributary weight carried by the bearings at 

different plan locations: one for the center column, one for the edge columns, and one for the 

corner columns. For the hybrid LR isolation system, we used the edge connection assembly for 

three of the four bearings. However, the center connection assembly was substituted at the East 

edge location, because erratic measurements were observed in the assembly used at that 

location in the prior TP configuration tests.

The plan drawings of the two connection assemblies utilized for the hybrid LR configuration are 

shown in Figure 4-1, and a photograph of a constructed assembly is shown in Figure 4-2. The 

center column assembly placed 3 Type A load cells on a circle 35 cm (13.8 in) from the center 

of both plates and 6 Type B load cells on a circle 90 cm (35.4 in) from the center of both plates, 

both with equal angular spacing. The edge column assemblies placed 1 Type A load cell at the 

center of both plates, and 6 Type B load cells on a circle 75 cm (29.5 in) from the center of both 

plates with equal angular spacing. The Type A and Type B load cells differed in their capacities 

as listed in Table 4-1. Drawings of the load cells are shown in Figure 4-3, which indicate the bolt 

pattern for the top and bottom rings and the elevation. The connection plates were produced by 

milling a steel plate with thickness = 10.2 cm (4 in) down to 9.5 cm (3.7 in), which leveled the 

surface. Because the two types of load cells differed in height, the thickness of each bottom 

connection plate was milled down to 9.1 cm (3.6 in) at Type A locations and 7.6 cm (3.0 in) at 

Type B locations (see Figure 4-4). The load cells were installed upside down between the top 

and bottom connection plates. The complete set of drawings for connection the load cell 

assemblies, LR bearings and testbed building to the simulator platform are given in Appendix C.
42

Figure 4-1 Load cell connection assemblies used for: (a) East bearing and (b) North, 

South and West bearings

Figure 4-2 Connection assembly

6 load cell B

3 load cell A

Plate connecting 

isolator to load cells

350 900

6 load cell B

Load cell A

Plate connecting 

load cells to table

Plate connecting 

isolator to load cells

750

(a)

(b)

Plate connecting 

load cells to table

Top 

connecting 

plate

Bottom 

connecting 

plate

Isolator

Load cell

Shake table
43

Table 4-1 Properties of load cells. 

Type Number Height Vertical 

capacity

Horizontal 

capacity

Vertical 

stiffness

Horizontal 

stiffness

(units) (cm) (kN) (kN) (kN/cm) (kN/cm)

A 44 18 400 250 85000 24000

B 32 19.5 700 400 140000 35000

Figure 4-3 Load cell drawings with bolt patterns and elevation views: (a) Type A and (b) 

Type B

(a) (b)
44

Figure 4-4 Connection of Type A and Type B load cells to top and bottom connection 

plates

4.2 Instrumentation

Approximately 470 channels were used for measuring the responses of structural and 

nonstructural components in the building with the hybrid LR isolation system. The sampling 

frequency of all channels was 1000 Hz. The results included in this report were based on 

measurements from the following three types of sensors:

• Sensors for measuring force: load cells (90 channels)

• Sensors for measuring displacement: displacement transducers (26 channels)

• Sensors for measuring acceleration: accelerometers (100 channels)

• The following describes the details of each sensor type.

Unless otherwise mentioned, all recorded data of the structural responses presented in the 

report was filtered using a Low-Pass Butterworth filter with a cut off frequency of 50 Hz. The 

filter shape as a function of normalized frequency is shown in Figure 4-5. The low pass filter 

“filtered out” or eliminated the high frequency components of the signal while preserving lower 

frequency components, including the dominant isolation frequency. The shape of the 

Butterworth filter provides a smooth transition from filtered to preserved frequencies. The cutoff 

frequency of 50 Hz was selected to eliminate noise that would not affect structural or equipment 

response and performance.
45

Figure 4-5 Magnitude of low-pass Butterworth Filter transfer function

4.2.1 Load Cells

Triaxial load cells were used to measure the forces in the LR bearings in all three directions. 

Figure 4-6 illustrates the load cell channels for each of the North, South, East and West 

bearings. The East bearing used a configuration of 9 load cells while the other bearings used 

only 7 load cells.

4.2.2 Displacement Transducers

Displacement transducers were used to measure the displacement of the isolation system and 

story drift. Figure 4-7 shows the layout of displacement transducers (wire pots) at base level for 

measuring the displacement of the isolation system. Three wire pots each were installed at the 

column bases at the North side and East side to measure the displacement in the x and ydirections, respectively. Three transducers were required to uniquely determine the translation 

and rotation of the isolation system and three additional channels were included for redundancy. 

Laser-based transducers were used to measure story drift. Each sensor was attached to a 

vertical instrumentation frame and its reflecting plate was attached to the floor above as shown 

in Figure 4-8. A pair of transducers measured the relative displacement between the two floors 

in each direction at 2 locations (Figure 4-9). Assuming a rigid floor diaphragm, 3 unique 

displacement transducers were needed for determining relative displacement between the 

adjacent floors. An additional displacement transducer was added in each story for redundancy.

The layout of the 4 displacement transducers in the 2nd to 5th story is shown Figure 4-9. In the 

1st story, the 4 displacement transducers were installed at the SE and NW columns.
46

 

Figure 4-6 Load cell channels for the hybrid LR isolation configuration

Y

X

E

S

N

W
47

Figure 4-7 Layout of displacement transducers at base

Figure 4-8 Instrumentation for measuring story drift

Displacement

transducer

Load 

cell

Isolator

Connecting 

plate

Y

X

N

Column 

base

Base 

girder

N

S

NE NW

E

SE SW

C W
48

Figure 4-9 Layout of displacement transducers to measure story drift in 2nd to 5th

stories

4.2.3 Accelerometers

Three triaxial accelerometers were installed to measure the 3 components of acceleration at the 

4 corners of the earthquake simulator platform. Accelerations at the center of the platform were 

also measured by permanent sensors integrated into the simulator control system. The 

measured acceleration at the center of the table included all 6 six components (3 translational 

components and 3 rotational components) of motion.

Two uniaxial accelerometers were also installed on the top plates of the connection assemblies 

(Section 4.1) to measure the horizontal acceleration of the plates (Figure 4-10). The recorded 

acceleration is used to derive the bearing forces. The load cells described in Section 4.2.1 did 

not measure the isolator force but rather the force just beneath the top connection plate. These 

forces differed by the inertia force associated with the top connection plate and the bottom half 

of the isolator. Since the total mass separating these two locations was large (about 4 tons) 

depending on location and the expected acceleration was also large (approximately equal to the 

input acceleration, about 1 g), the inertia force, which was significant compared to the isolator 

force, was accounted for.

Laser transducer

Support 

truss
49

Figure 4-10 Accelerometers at the top connection plates to estimate inertia forces

Floor accelerations (2 horizontal and vertical components) were measured using 3 triaxial 

accelerometers installed at the SE, NE and NW corners of every floor. These triaxial 

accelerometers were attached to the column face just above the floor slab. Vertical 

accelerations at other locations on the floor slab were also recorded. Figure 4-11 shows the 

layout of accelerometers on the 5th floor, which was a typical layout for all floors. The vertical 

accelerometers were attached to the bottoms of the slabs. 

Figure 4-11 Layout of accelerometers at the 5th floor

Acceleromete

Isolator

Load cell

Bottom connecting plate

Shaking table

Top connecting plate

Triaxial 

accelerometer

Vertical 

accelerometer
50

4.3 Installation of the Specimen on the Earthquake Simulator

The connection assemblies were put together using the following process. Holes were drilled 

and tapped in the steel connection plates as needed. The load cells were first bolted to the top 

hexagonal shaped plate of the connection assembly (Figure 4-2). The bottom plate was then 

added to the assembly (Figure 4-12). The connection assemblies were then turned over and 

bolted to the earthquake simulator platform using 48 mm (1.9 in) diameter threaded rods (see 

Figure 4-13). The installation of the testbed building with the hybrid LR isolation system, 

immediately followed testing and removal of the TP isolation system. After removal of the TP 

bearings and rearrangement of the connection assemblies, the LR bearings were bolted to the 

connection assemblies while CL bearings were bolted directly to the simulator platform (Figure

4-14). The building was then transported across the laboratory using two 400-ton-cranes and 

lowered over the isolation system (Figure 4-15). 

The testbed building was connected to the isolation system through bolt holes in the column 

bases that had been drilled and tapped from below (Figure 4-16). Drilling and tapping these 

holes in the overhead position was a laborious and expensive process (Figure 4-17). As such, 

measures were taken to limit the number and size of connecting bolts. Four M24 bolts were 

used to connect each LR bearing to the structure above and the connecting plates below. These 

bolts were attached through 30 mm (1.2 in) oversized holes in the top connecting plates and 33 

mm (1.3 in) oversized holes in the bottom connecting plate. The oversized holes were used to 

help align the testbed building with the isolators. 

Figure 4-12 Connection assembly
51

Figure 4-13 Connection assemblies on the simulator platform

Figure 4-14 Installing the LR bearings to the connecting assemblies and the CL 

bearings to the simulator platform
52

Figure 4-15 Lowering the testbed building over the isolation system

Figure 4-16 Bolt holes for connecting the testbed structure to the isolation system

Bottom face of 

the column base

Bolt hole for 

connecting the 

specimen to 

the LR 

bearings
53

Figure 4-17 Drilling and tapping holes at the bottom of the specimen

The measured weight of the building (excluding the isolators) was 5220 kN (1173 kip). This 

weight was determined during the testing of the TP isolation configuration, as described in Dao 

(2012), since static forces were measured in every bearing. The measured weight of the testbed 

was about 17% greater than the 4540 kN (1020 kips) anticipated in the design (Section 3.2). 

The change in weight affected the realized stiffness and strength of the isolation system. 

However, since the properties of the LR bearings are highly amplitude dependent (see Section 

7.4), the influence of this change in supported weight was not explored in detail.

The static vertical load on an isolator is expected to be proportional to the mass of the tributary 

area. This condition could have been obtained if the testbed had been erected directly on the 

isolation system (similar to the expected construction process). However, the testbed had been 

built and stored outside for more than 2 years before testing so that its base was warped and 

the distribution of vertical load on all isolators deviated from the calculated values. Table 4-2 

shows the measured vertical load on each LR bearing after the building was bolted to the 

isolation system and the expected vertical load on the isolators based on the pre-test simulation 

model. The expected vertical load on each isolator was scaled by 1.17 compared to the values 

reported earlier (Figure 3-7) to adjust for the actual weight (=5220 kN or 1174 kips) of the 

testbed. The table shows that most of the LR bearings were carrying less gravity load than 

expected based on tributary area. The static load in the West bearing was very small at 235 kN 

(53 kip). Shims were installed to achieve a distribution of gravity load similar to that assumed for 

design, with limited success. 

Table 4-2 Vertical load on each bearing after installation. 

Bearing E S N W Percent 

of Total

Actual load (kN) 435 755 490 235 37%

Expected load (kN) (*) 749 708 725 468 51%

Difference (%) -41.9 6.6 -32.4 -49.9 

(*) The expected load was linearly scaled such that the total vertical load matches the 

measured load
54

4.4 Test Schedule

The test schedule included 3 days of shaking (21 simulations) for the TP isolation configuration, 

2 days of shaking (15 simulations) for the hybrid LR configuration, and 1 day of shaking (5 

simulations along with white noise and sine sweep) for the fixed-base configuration. Dao (2012) 

discussed the selection of the ground motions for the TP configuration, while Chapter 3 

described the selection of ground motions for the hybrid LR configuration. Since only a minimal 

number of simulations could be included for the fixed-base configuration, motions were selected 

that would allow a broad comparison between the three different configurations, and to provide 

insight into the influence of vertical excitation. 

Although the majority of the inputs were 3D, some of the earthquake motions were applied as 

horizontal (XY) only, which enables a study of the effect of vertical excitation. Other reasons for 

not including the vertical component of excitation for some simulations included: 1) lack of 

access to the vertical record, 2) the capacity of the earthquake simulator limited the application 

of all 3 components at full scale, and 3) test-day decisions to limit damage to nonstructural 

components. 

For the isolation configurations, the floors containing nonstructural components and contents 

were inspected only at the end of the test day, with one exception. On the first test day (TP 

isolation configuration), nonstructural components and contents were inspected after Rinaldi 

88%, which generated some unexpected ceiling damage and disruption to contents. The 

shaking of the fixed-base building, completed in 1 day, included 5 earthquake excitations. For 

the fixed-base configuration only, nonstructural components were inspected and partially 

repaired after every simulation, thus 3D white noise excitation preceded and followed every 

earthquake simulation for system identification before and after the repairs. Unidirectional white 

noise excitations were also applied at the beginning and end of the day. Damage to 

nonstructural components and content disruption was observed in all system configurations 

under large intensity vertical excitation, and the damage was similar in all configurations. The 

nature of the damage is not discussed in this report, as the tested nonstructural components 

were not nuclear qualified, and the observations could be misleading. Rather the performance 

of safety related nuclear structures should be judged on the basis of the reported seismic 

demand parameters. Further information about the response of the nonstructural components is 

provided in Soroushian et al. (2012).

For completeness, the simulation schedule for all three building configurations is listed in Tables 

4-3 to 4-5. In Tables 4-3 and 4-4, the shaded simulations are 3D excitations and the rest are XY 

simulations. Color coded groups of simulations represent the same input at different scale 

factors; for instance, red = the Vogtle suite of motions in Table 4-4. In the shaking schedule for 

the fixed-base building (Table 4-5), the earthquake simulations are shaded and the motion in 

red use the same input with different scale factors. The nomenclature introduced in Table 4-4 

(simulation abbreviation) is used throughout the report to refer to different simulations in the 

hybrid LR system. The abbreviation consists of the first three letters of the station name with the 

scale factor. If the input excitation is not 3D, then “(XY)” is added for bidirectional horizontal 

input and “(Y)” is added for unidirectional input in the y-direction. If the simulation is repeated 

with the same input, the repetitions are labeled “-1” and “-2”. Thus, “SIN100(Y)-1”reflects the 

first repetition of a sine wave input scaled to 100% and the input is unidirectional in the ydirection; “VOG150” reflects the Vogtle record scaled to 150% and the input is 3D.
55

Table 4-3 Simulation schedule for the TP isolation configuration. 

Date

dd/mm/y

y

Time

Simulation

Abbreviatio

n

Motion

Scale Factor Damage

Inspection X Y Z 

17/08/11

12:01:46 SIN65(X) Sine-wave 0.65 0 0 

12:49:54 SIN100(X) Sine-wave 1.00 0 0 

13:42:20 WSM80

Superstition 

Hills, 

Westmorland,

0.80 0.80 0.80

14:30:21 ELC130

Imperial Valley,

El Centro

1.30 1.30 1.30

15:20:16 RRS88

Northridge,

Rinaldi Rec. Sta.

0.88 0.88 0.88 Yes

17:16:16 SYL100 Northridge, 

Sylmar 1.00 1.00 1.00

17:48:56 TAB50 Tabas, Tabas 

Sta. 0.50 0.50 0.50 Yes

18/08/11

11:35:31 LGP70

Loma Prieta

Los Gatos Pres. 

Ctr.

0.70 0.70 0.70

12:25:40 TCU50(XY) ChiChi, TCU065 0.50 0.50 0 

13:55:30 TCU70(XY) ChiChi, TCU065 0.70 0.70 0 

14:31:59 IWA100(XY) Tohoku, 

Iwanuma 1.00 1.00 0 

15:45:46 SAN100(XY) Sannomaru 1.00 1.00 0 

16:34:58 TAK100 Kobe, JMA 

Takatori 1.00 1.00 1.00

17:05:03 KJM100 Kobe, Kobe JMA 1.00 1.00 1.00 Yes

19/08/11

11:29:55 RRS88(XY)

Northridge,

Rinaldi Rec. Sta.

0.88 0.88 0 

12:16:55 TCU80(XY) ChiChi, TCU065 0.80 0.80 0 

13:08:07 TAB80 Tabas, Tabas 

Sta. 0.80 0.80 0.80

14:02:19 TAB90(XY) Tabas, Tabas 

Sta. 0.90 0.90 0 

14:50:46 TAB100(XY) Tabas, Tabas 

Sta. 1.00 1.00 0 

15:28:19 SCT100(XY) Michoacan, SCT 1.00 1.00 0 

16:19:03 TAK115 Kobe, JMA 

Takatori 1.15 1.15 1.00 Yes
56

Table 4-4 Simulation schedule for the hybrid LR isolation configuration. 

Date

dd/mm/y

y 

Time

Simulation

Abbreviatio

n 

Motion

Scale Factor Damage

Inspection X Y Z 

25/08/11

11:19:52 WSM80 Superstition Hills, 

Westmorland 0.80 0.80 0.80

12:21:52 SIN100(Y)-1 Sine-wave 0 1.00 0 

13:06:04 VOG75-1 Vogtle #13 0.75 0.75 0.75

13:56:09 VOG100 Vogtle #13 1.00 1.00 1.00

14:33:53 VOG125 Vogtle #13 1.25 1.25 1.25

15:15:09 VOG150 Vogtle #13 1.50 1.50 1.50

16:17:50 VOG175 Vogtle #13 1.75 1.75 1.75

16:52:49 DIA80 Diablo #15 0.80 0.80 0.80 Yes

26/08/11

12:03:09 DIA95(XY) Diablo #15 0.95 0.95 0 

12:48:49 ELC130 Imperial Valley, 

El Centro 1.30 1.30 1.30

13:44:36 IWA100(XY) Tohoku, Iwanuma 1.00 1.00 0 

14:37:30 RRS88(XY) Northridge

Rinaldi Rec. Sta. 0.88 0.88 0 

15:20:52 RRS88 Northridge

Rinaldi Rec. Sta. 0.88 0.88 0.88

16:15:12 VOG75-2 Vogtle #13 0.75 0.75 0.75

16:59:19 SIN100(Y)-2 Sine-wave 0 1.00 0 Yes
57

Table 4-5 Simulation schedule for the fixed-base building. 

Date

(dd/mm/yy) Time Simulation abbreviation Motion

Scale factor Damage 

inspection X Y Z 

31/08/11

10:19:52 WHT100(X)-1 White noise 1.00 0 0 

10:30:02 WHT100(Y)-1 White noise 0 1.00 1.00 

10:38:32 WHT100(Z)-1 White noise 1.00 1.00 1.00 

10:50:35 WSM80

Superstition 

Hills, 

Westmorland

0.80 0.80 0.80

11:02:50 WHT100-1 White noise 1.00 1.00 1.00 Yes

12:06:31 WHT100-2 White noise 1.00 1.00 1.00 

12:18:47 RRS35(XY)

Northridge, 

Rinaldi Rec. 

Sta.

0.35 0.35 0 

12:28:02 WHT100-3 White noise 1.00 1.00 1.00 Yes

13:37:34 WHT100-4 White noise 1.00 1.00 1.00 

13:51:20 RRS35

Northridge, 

Rinaldi Rec. 

Sta.

0.35 0.35 0.35

14:03:01 WHT100-5 White noise 1.00 1.00 1.00 Yes

15:12:50 WHT100-6 White noise 1.00 1.00 1.00 

15:24:53 RRS35(XY)88(Z) 

Northridge,

Rinaldi Rec. 

Sta.

0.35 0.35 0.88

15:33:51 WHT100-7 White noise 1.00 1.00 1.00 Yes

17:07:04 WHT100-8 White noise 1.00 1.00 1.00 

17:22:33 IWA70(XY) Tohoku, 

Iwanuma 0.70 0.70 0 

17:35:28 WHT100(X)-2 White noise 1.00 0 0 

17:43:12 WHT100(Y)-2 White noise 0 1.00 0 

17:52:47 WHT100(Z)-2 White noise 1.00 1.00 1.00 Yes
58

4.5 Table Motions

The peak accelerations of the target motions and the actual motions generated by the 

earthquake simulator are compared in Tables 4-6 and 4-7. The target records were obtained 

from various sources such as the PEER NGA database (Chiou et al. 2008), Huang et al. (2009), 

and sources within E-Defense; and scaled by the scaled factors given in Tables 4-3 to 4-5. The 

realized input motions to the structure were generally amplified by the earthquake simulator 

relative to the target motions, and amplification factors of 50% were not uncommon. 

Amplification occurred because the recommended iterative response modification technique, 

which involves gradually increasing the intensity of the motions while making modifications to 

the control settings, was not used in favor of performing more simulations with a wider variety of 

earthquakes. We observed, however, that ground motion excitations were replicated 

consistently when repeated for the different system configurations. 

Amplification was notable during the Northridge-Rinaldi (RRS88) simulation due to its effect on 

the vertical response of the system. The motion was reproduced similarly for all three building 

configurations. The acceleration histories of the 3 components of this motion for the hybrid LR

configuration are plotted in Figure 4-18. The horizontal amplification occurred at the instant of 

the large horizontal pulse in the record and the vertical acceleration was amplified at the same 

instant. Thus, the realized intensity of the Rinaldi motion was much stronger than the intended 

design level earthquake, and in particular the vertical component of excitation might be 

considered extreme. On the other hand, motions with similar vertical intensity can be found in 

the PEER NGA database (Chiou et al. 2008). Note that the realized intensity of vertical 

excitation in RRS88(XY) was non-negligible (vertical peak ground acceleration or PGA = 0.05g

for the hybrid LR configuration and 0.10g for TP configuration).

The 5% damped response spectra are compared for the target motions and the realized input 

motions. The ratio between these spectra at periods ranging from 0.01 sec to 5 sec is plotted in 

Figures 4-19 and 4-20. At periods longer than 0.7 sec, the spectral amplitudes of the realized 

motions did not differ much from the target motions. At periods less than 0.7 sec, the spectral 

amplitude of the realized motions in the horizontal direction was generally larger than that of the 

target motions. The isolation system is controlled by the post-yield properties of the bearings, so 

that the isolator displacement demands would not be significantly affected by the difference 

between the realized motions and the target motions. However, the contribution of higher mode 

effects to structural accelerations may have been amplified in the simulations compared to 

typical ground motions. The earthquake simulator appeared to amplify the horizontal period 

components at around 0.2 sec more strongly than components at neighboring periods.
59

Table 4-6 Peak acceleration of target and realized motions for the TP isolation 

configuration. 

Trial

Peak XX (gg) Peak YY (gg) Peak ZZ (gg)

Target Table Target Table Target Table

TP Configuration

80WSM 0.171 0.169 0.135 0.147 0.174 0.140

130ELC 0.278 0.293 0.408 0.484 0.263 0.261

88RRS 0.427 0.586 0.730 1.213 0.722 1.241

100SYL 0.601 0.674 0.869 1.145 0.519 0.543

50TAB 0.450 0.585 0.418 0.463 0.327 0.357

70LGP 0.415 0.445 0.391 0.628 0.641 0.687

50TCU 0.408 0.453 0.304 0.278 0.000 0.015

70TCU 0.571 0.648 0.425 0.378 0.000 0.027

100IWA 0.364 0.409 0.418 0.580 0.000 0.031

100SAN 0.190 0.231 0.167 0.161 0.000 0.020

100TAK 0.747 0.789 0.619 0.922 0.288 0.259

100KJM 0.595 0.680 0.822 0.893 0.340 0.408

88RRSXY 0.427 0.532 0.730 1.194 0.000 0.098

80TCU 0.653 0.747 0.486 0.418 0.000 0.034

80TAB 0.720 0.870 0.670 0.836 0.523 0.593

90TAB 0.810 0.930 0.753 1.011 0.000 0.102

100TAB 0.901 0.995 0.837 1.139 0.000 0.120

100SCT 0.171 0.177 0.101 0.106 0.000 0.017

115TAK 0.859 0.936 0.712 1.088 0.288 0.278
60

Table 4-7 Peak acceleration of target and realized motions for hybrid LR isolation and 

fixed-base configurations. 

Trial

Peak XX (gg) Peak YY (gg) Peak ZZ (gg)

Target Table Target Table Target Table

Hybrid Configuration

WSM80 0.170 0.195 0.138 0.150 0.209 0.145

VOG75-1 0.329 0.391 0.213 0.259 0.215 0.214

VOG100 0.438 0.521 0.284 0.346 0.286 0.297

VOG125 0.548 0.687 0.355 0.451 0.358 0.368

VOG150 0.657 0.857 0.426 0.549 0.429 0.437

VOG175 0.767 1.025 0.497 0.639 0.501 0.493

DIA80 0.783 0.917 0.543 0.662 0.455 0.452

DIA95(XY) 0.930 1.118 0.645 0.808 0.000 0.063

ELC130 0.278 0.300 0.406 0.497 0.259 0.277

IWA100(XY) 0.363 0.429 0.420 0.590 0.000 0.021

RRS88(XY) 0.430 0.524 0.733 1.180 0.000 0.051

RRS88 0.430 0.521 0.733 1.193 0.738 1.257

VOG75 0.329 0.393 0.213 0.246 0.215 0.220

Fixed-base 

configuration

80WSM 0.171 0.219 0.135 0.175 0.174 0.136

35RRSXY 0.170 0.201 0.290 0.398 0.000 0.011

35RRS 0.170 0.201 0.290 0.406 0.287 0.350

88RRS 0.170 0.228 0.290 0.409 0.722 1.062

70IWA 0.255 0.270 0.292 0.373 0.000 0.013
61

Figure 4-18 Acceleration history of target and realized RRS88 motion in the hybrid LR 

isolation configuration test

-1.5

-1

-0.5

0

0.5

1

1.5

Acceleration ax (g)

-1.5

-1

-0.5

0

0.5

1

1.5

Acceleration ay (g)

5 10 15 20 25

-1.5

-1

-0.5

0

0.5

1

1.5

Acceleration az (g)

Time (sec)

Table motion

Target motion
62

Figure 4-19 Ratio of realized to target motion 5% damped spectral accelerations – hybrid 

LR configuration

0

0.5

1

1.5

SAxTable/SAxTarget

0

0.5

1

1.5

SAyTable/SAyTarget

10-2 10-1 100 0

0.5

1

1.5

SAzTable/SAzTarget

Period (sec)
63

Figure 4-20 Ratio of realized to target motion 5% damped spectral accelerations – fixedbase configuration

10-2 10-1 100 101 0

0.5

1

1.5

2

SAxTable/SAxTarget

10-2 10-1 100 101 0

0.5

1

1.5

2

SAyTable/SAyTarget

10-2 10-1 100 101 0

0.5

1

1.5

2

Period, T (s)

SAzTable/SAzTarget
64

4.6 Derived Responses

4.6.1 Horizontal Displacement of the Isolation System

An algorithm to compute the displacements in each isolator from the measured displacements 

in the string pots is described next. The algorithm accounts for large displacement geometric 

effects as a result of the large displacement demand in the bearings. From the original and 

displaced configurations of the isolation system in Figure 4-21, the coordinates XXAA′ , YYAA′ ... of 

displaced nodes A’, B’, D’, F’, G’ and H’ are:

XXAA′ = ΔΔ + XXAA cccccc φφ − YYAA ssss φφ

YYAA′ = ΔΔ + XXAA ssssφφ + YYAA cccccc φφ

XXBB′ = ΔΔ + XXBB cccccc φφ − YYBB ssss φφ

YYBB′ = ΔΔ + XXBB ssss φφ + YYBB cccccc φφ

XXDD′ = ΔΔ + XXDD cccccc φφ − YYDD ssss φφ

YYDD′ = ΔΔ + XXDD ssss φφ +

YYDD cccccc φφ

(4.1)

XXFF′ = ΔΔ + XXFF cccccc φφ − YYFF ssss φφ

YYFF′ = ΔΔ + XXFF ssss φφ + YYFF cccccc φφ

XXGG′ = ΔΔ + XXGG cccccc φφ − YYGG ssss φφ

YYGG′ = ΔΔ + XXGG ssss φφ + YYGG cccccc φφ

XXHH′ = ΔΔ + XXHH cccccc φφ − YYHH ssss φφ

YYHH′ = ΔΔ + XXHH ssssφφ + YYHH cccccc φφ

where ΔXX, ΔYYand φφ are the displacements and rotation at the center bearing in moving from the 

original configuration C to the displaced configuration C’, with sign convention shown in Figure 

4-21(b); XXAA, YYAA are coordinates of the original point A, and so on.

From the displaced configuration in Figure 4-21(b):

(XXAA′ − XXaa)2 + (YYAA′ − YYaa)2 = LLAA

2

(XXBB′ − XXbb)2 + (YYBB′ − YYbb)2 = LLBB

2

(XXDD′ − XXdd)2 + (YYDD′ − YYdd)2 = LLDD

2

(4.2)

�XXFF′ − XXff�

2

+ �YYFF′ − YYff�

2 = LLFF

2

�XXGG′ − XXgg�

2

+ �YYGG′ − YYgg�

2 = LLGG

2

(XXHH′ − XXh)2 + (YYHH′ − YYh)2 = LLHH

2
65

where XXaa, YYaa are coordinate of node a; LLAA = LLAA0 + ΔLLAA is the distance between a and A’; LLAA0 is 

the original length of the transducer and ΔLLAA is the change in length measured by the 

transducer.

Substituting Equation (4.1) into Equation (4.2) leads to a system of 6 nonlinear equations to 

solve for 3 unknown ΔXX, ΔYY and φφ. The system of equations was solved using the lsqnonlin 

command in Matlab, which is applicable to nonlinear least-squares (nonlinear data-fitting) 

problems. After solving for ΔXX, ΔYY and φφ, the coordinate of the displaced isolators were 

determined from Equation (4.1). These displacements were determined by subtracting the 

original coordinates from the displaced coordinates. 

f g X h 

SE S

E C W

N

SW

NE NW

a

b

d

f

g

h

A

B

D

F G H

A’

B’

D’

F’

G’

H’

X

Y

LLYY

LLXX

C’

C

C’

ΔYY

ΔXX φ

a

b

d

Y

LLAA

LLBB

LLDD

LLFF

LLGG

LLHH

a)

b)
66

Figure 4-21 Configurations for solving displacement of the isolation system. (a) original 

configuration, (b) displaced configuration
67

4.6.2 Isolator Forces

The X, Y and Z components of the recorded dynamic force from all load cells of an isolator were 

added to get the X, Y and Z components of the dynamic reaction at the load cells level. This 

reaction was then modified by the inertia forces of the connection plate and the bottom concave 

plate of the bearing to get the dynamic reaction at the isolator level. From the free body diagram 

in Figure 4-22 the relationships between the dynamic reaction components RRXX, RRZZ at the isolator 

level and the dynamic reaction components RRcc ,RRcc at the load cell level are:

X cX c cX

Y cY c cY

Z cZ c cZ

R R ma

R R ma

R R ma

= −

= −

= −

(4.3)

where mmcc is the mass of the top plate in the bearing connection assemblies plus the bottom half 

of the bearing; and aacc , aacc and aacc represent the horizontal and vertical accelerations recorded 

at the top connection plate. The reactions RRXX and RRZZ in Equation (4.3) represent the forces at 

mid-height of the LR bearing. These reactions are dynamic reactions so that the participation of 

the gravity load is not included in the equations.

Since vertical acceleration in the top connection plate was not recorded, the vertical 

acceleration in the earthquake simulator was used for aacc , which approximates the load cells as 

vertically rigid. The vertical acceleration of the simulator platform at every isolator was 

extrapolated from the measured acceleration at the center of the platform including the effect of 

roll and pitch components. The validity of these extrapolation accelerations was checked by 

comparing the extrapolated acceleration at the 4 corners of the platform to the accelerations 

recorded at these locations. 

Figure 4-22 Free body diagram illustrating derivation of isolator reaction

Colum base

Top plate of the LR 

bearing

Bottom plate of the LR 

bearing

Top connecting 

plate

RRXX

RRXX

RRcc

aacc

RRcc

mmccaacc

RRZZ

RRZZ aacc

mmccaacc
68

The vertical force in all load cells was measured before each test series. The load cells were 

then zeroed before the first simulation of the test series so that only the dynamic force variation 

was measured during the simulations. The forces were only sampled during the simulation so 

that any redistribution of forces on the bearings from the original static state were reflected as 

offsets in the vertical forces at the beginning of each new simulation. The procedure used to 

measure the initial static forces in the LR bearings was found to be unreliable; thus, the 

computed static loads may have errors in them. Fortunately, interpretation of the LR bearing 

response was not sensitive to the measured vertical force.

4.6.3 Horizontal Acceleration and Story Drifts

As shown in Figure 4-11, the horizontal accelerations were measured at the SE, NW and NE 

corners of each floor. These recorded accelerations were processed to get an average 

acceleration in each direction, computed as: 

( )

( )

,

,

1

3

1

3

x avg xSE xNE xNW

y avg ySE yNE yNW

a aaa

a aaa

= ++

= ++

(4.4)

where aaxxxxxx, aayy are XX and YY −components of the horizontal acceleration at the SE corner, and 

so on. Physically, the average acceleration represents a plan location one third of the way from 

the geometric center to the NE corner of the building.

The story drift in X- and Y-direction at the geometric center were also interpolated from the 

measured story drift at the 2 locations shown in Figure 4-23. For instance, the story drifts δδxxxx at 

(Figure 4-23) or δδyy the geometric center were extrapolated from the story drifts at the SE and 

NW corners δδxxxxxx, δδxxxxxx , δδyy , δδyy as follows:

( )

( )

1

2

1

2

xC xSE xNW xSE

yC ySE yNW ySE

L

L

L

L

δδ δ δ

δδ δ δ

=+ −

=+ −

(4.5)

Inconsistencies were observed in the drift sensor measurements, especially under vertical 

excitation. The vertical slab vibration is believed to have produced rocking of the measurement 

towers, which distorted the recorded drifts.
69

Figure 4-23 Diagram illustrating the computation of drift at the geometric center

Laser transducer

Geometric 

center

Y

X

LL1

LL2

δδxxxxxx

δδxxxxxx

δδxxxx

71

5. SUMMARY OF EXPERIMENTAL SIMULATION RESULTS FOR THE 

HYBRID LR ISOLATED BUILDING

This chapter summarizes the overall response of the building with hybrid LR isolation system, 

with emphasis on the peak demands of various response parameters observed throughout the 

experimental program. Response quantities examined include displacement, rotation, shear 

force, axial force in compression and tension, and residual displacement of the isolators; and 

floor accelerations and story drifts in the structure.

5.1 Isolator Displacements

The target isolator displacements were 30 cm (11.8 in) for the design base earthquake (DBE) 

and 60 cm (23.6 in) for the maximum credible earthquake (MCEbeyond DBE). However, as 

mentioned previously, the maximum isolator displacement targeted in the test program was 55 

cm (21.6 in) due to the various safety-related limitations imposed by the test facility (see Section 

3.3.1). The DBE motion, VOG100, produced a peak vector sum displacement of 26.5 cm (10.4 

in) and the MCE beyond DBE motion, VOG175, produced a peak vector sum displacement of 

50.5 cm (19.9 in) as shown in Figure 5-1. The peak displacements observed during the Vogtle 

suite of simulations were slightly lower than analytically predicted, and did not reach the target. 

However, the displacement demands for the Diablo Canyon suite of simulations were slightly 

greater than predicted. The scale factor for the largest Diablo Canyon simulation (DIA95(XY)), 

originally planned for 100%, was adjusted on the day of testing to achieve the target 

displacement of 55 cm (21.6 in). The peak displacement observed in any LR bearing during 

DIA95(XY) was 54.7 cm (21.5 in). The smallest displacement (8.8 cm or 3.5 in) was observed 

during the service level simulation WSM80. The peak displacement increased approximately 

linearly as the simulation intensity was increased from VOG75 to VOG175. Because of the 

nonlinearity of the isolation system, the displacement demand would not normally be expected 

to increase linearly with excitation intensity.

The maximum displacements observed in each LR bearing (East (E), South (S), North (N) and 

West (W)) are summarized in Figure 5-2 for the x-direction, y-direction, and overall peak in any 

direction, determined as the peak of the vector sum displacement history. The simulations 

names are abbreviated by numbers in the figure, in order of their sequence, where the 

correspondence between number and simulation name, the directions that the excitations were 

applied and the input scale factor is summarized in Table 5-1 for convenience. By way of the 

small rotation assumption used to process the sensor data and derive individual isolator 

displacements (Section 4.6.1), the x-direction displacements were identical for the North and 

South bearings, which had the same y-coordinate, and the y-direction displacements were 

identical for the East and West bearings, which had the same x-coordinate. The East bearing 

experienced the largest displacement for most of the simulations (Figure 5-2(c)) due to the 

observed base rotation (see Section 5.2). The displacement traces (displacement in x-direction 

versus displacement in y-direction) of the East LRB are compared for four simulations in Figure 

5-3: (a) WSM80, which produced the smallest displacement demand, (b) DIA95(XY), which 

produced the largest displacement demand, (c) VOG100, which was scaled to DBE intensity, 

and (d) VOG175, which was scaled to MCE intensity. The displacement observed in WSM80 

was trivially small compared to the other simulations, and the large discrepancy in 

displacements affected the ability to model the LR bearings with a single set of physical 

parameters (discussed in Chapter 7). The simulations produced both linear and circular 

displacement orbits in the bearings, the latter of which would be more affected by bidirectional 

coupling. As mentioned in Section 3.3.2, the Vogtle ground excitation was rotated 11.25 

degrees to induce the maximum displacement in a diagonal direction, as observed in Figure 5-

3(d).
72

Figure 5-1 Peak vector sum displacement recorded in any LR bearing for each 

earthquake simulation

Figure 5-2 X-direction, y-direction and overall peak (vector sum) displacement of each 

LR bearing for each earthquake simulation

0

10

20

30

40

50

60

WSM80

SIN100(Y)-1

VOG75-1

VOG100VOG125VOG150VOG175

DIA80

DIA95(XY)

ELC130

IWA100(XY)

RRS88(XY)

RRS88

VOG75-2

SIN100(Y)-2

Peak Vector Displacement (cm)

8.8

21.1 19.6

26.5

34.5

42.3

50.5

42.5

54.7

20.6

37.7 38.8 38.5

21 22.2

0

20

40

60

X-direction

0

20

40

60

Y-direction

Peak Displacement (cm)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

0

20

40

60

Vector sum

GM #

E S N W
73

Table 5-1 List of ground motion (GM) simulations by number, name, abbreviation and the 

directions considered.

GM # GM Name GM 

Abbreviation

Input 

Direction

Scale Factor 

(%)

1 Superstition Hills, 

Westmorland WSM80 X, Y, Z 80

2 Sine-wave SIN100(Y)-1 Y 100

3 Vogtle #13 VOG75-1 X, Y, Z 75

4 Vogtle #13 VOG100 X, Y, Z 100

5 Vogtle #13 VOG125 X, Y, Z 125

6 Vogtle #13 VOG150 X, Y, Z 150

7 Vogtle #13 VOG175 X, Y, Z 175

8 Diablo #15 DIA80 X, Y, Z 80

9 Diablo #15 DIA95(XY) X, Y 95

10 Imperial Valley, El Centro ELC130 X, Y, Z 130

11 Tohoku, Iwanuma IWA100 X, Y, Z 100

12 Northridge, Rinaldi Rec. Sta. RRS88(XY) X, Y 88

13 Northridge, Rinaldi Rec. Sta. RRS88 X, Y, Z 88

14 Vogtle #13 VOG75-2 X, Y, Z 75

15 Sine-wave SIN100(Y)-2 Y 100

Figure 5-3 Displacement trace (x vs y-direction displacement) of the East LRB for (a) 

WSM80, (b) DIA95(XY), (c) VOG100 and (d) VOG175

-50 0 50

-60

-40

-20

0

20

40

60

8.8

(a) WSM80

-X Disp (cm)

Y Disp (cm)

-50 0 50

-60

-40

-20

0

20

40

60

54.7

(b) DIA95 (XY)

-X Disp (cm)

Y Disp (cm)

-50 0 50

-60

-40

-20

0

20

40

60

26.5

(c) VOG100

-X Disp (cm)

Y Disp (cm)

-50 0 50

-60

-40

-20

0

20

40

60

50.5

(d) VOG175

-X Disp (cm)

Y Disp (cm)
74

5.2 Torsional Response

The dynamic characteristics of the testbed building were affected by stiffness asymmetry 

resulting from the unequal bays widths (equal to 7 m or 23 ft and 5 m or 16.4 ft) in the ydirection (Section 2.1, Figure 2-2), and various sources of mass eccentricity, the most notable 

being the asymmetrically configured steel blocks at the roof level (Section 2.3). The level of 

eccentricity is later quantified while discussing the model development for numerical simulation 

(Section 8.3). Aside from the supplementary roof weight, the sources of eccentricity were mild 

and typical of practice. However, as discussed in Section 3.3.2, the restrictions on the 

experimental setup did not allow for the isolation system to be configured to minimize torsion, 

unlike the design of a realistic structure with hundreds of isolators. Thus, non-negligible rotation 

was observed in the hybrid LR isolation system. 

The peak rotation angle at the base (isolation system) level for each simulation is summarized 

in Figure 5-4. The rotation angle observed during the sine wave simulation (SIN100-1 and 

SIN100-2) was negligible, since the sine wave was applied unidirectionally in the y-direction 

(theoretically uncoupled) to minimize the torsional response for bearing characterization. During 

WSM 80%, which produced the smallest displacement demand, a peak rotation angle of 0.0019 

rad was observed, and during VOG175 the largest peak rotation angle of 0.019 rad was 

observed. The peak rotation was not necessarily proportional to the peak displacement; for 

example, the greatest individual bearing displacement was observed in DIA95(XY) (Figure 5-1) 

but the greatest rotation angle was observed in VOG175 (Figure 5-4). The influence of this 

rotation on the bearing displacements can be observed from the displacement traces of all 

bearings during the VOG175 motion (Figure 5-5). From the SE to the NW isolator, the 

displacement traces transitioned from nearly linear (back and forth) motion to a circular 

displacement orbit. Furthermore, the peak displacement in the LR bearings varied from 50.5 cm 

in LRB-E to 44.6 cm in LRB-W, which is a 13% variation across the plan. Considering all 

isolation devices, the peak displacement varied from 53.1 cm (NE corner) to 42.7 cm (SW 

corner), a 24% variation from corner to corner. For the Rinaldi simulation that was repeated at 

the same scale factor for XY and 3D input, the peak rotation increased 7% from RRS88(XY) to 

RRS88 (Figure 5-4) while the peak displacement remained about the same, which may have 

been related to a residual rotation or displacement. 

Figure 5-4 Peak rotation angle of the base for each earthquake simulation

0

0.005

0.01

0.015

0.02

WSM80

SIN100(Y)-1

VOG75-1

VOG100VOG125VOG150VOG175

DIA80

DIA95(XY)

ELC130

IWA100(XY)

RRS88(XY)

RRS88

VOG75-2

SIN100(Y)-2

Peak Rotation (rad)
75

Figure 5-5 Displacement trace of each isolator during the Vogtle 175% simulation

5.3 System Base Shear

The total base shear was computed by summing the recorded shear force of the four LR 

bearings in the x and y-directions, evaluating the vector sum of the x and y-components, and 

determining the peak over all times steps. This procedure was applied in the two horizontal 

directions, and from this the peak vector sum was determined. The calculation of the base shear 

does not include forces in the CL bearings, which were not recorded in this experiment. The 

forces in the CL bearings were assumed to be negligible as the rated friction coefficient was 

about 0.0025. Although we cannot be certain of the influence of the CL bearings on the base 

shear, there was no evidence to suggest that the forces in the CL bearings were significant. The 

peak values of total (vector sum), x and y-direction base shears are listed in Table 5-2 for 

each simulation. The corresponding values of normalized base shear or base shear 

coefficient, listed in Table 5-3, were obtained by dividing the total base shear by the total 

static weight of the building. Figure 5-6 illustrates the results of Table 5-2 graphically. The 

largest base shear of 1467 kN (328 kips), corresponding to a base shear coefficient of 0.28, was 

observed during VOG175. Among the other largest base shear coefficients observed were 

DIA95(XY) (0.28), DIA80 (0.24) and RRS88(XY) and RRS88 (both 0.26).

-50

0

50 48.2

SE

44.7

S

Vogtle 175%

42.7

SW

-50

0

50 50.5

E

-X displacement (cm)

46.9

C

44.6

W

-50 0 50

-50

0

50 53.1

NE

-50 0 50

49.4

N

Y displacement (cm)

-50 0 50

47

NW
76

Table 5-2 Peak base shear for all simulations: total, x and y directions. 

GM # GM Name Peak Base 

Shear (kN)

Peak Base 

Shear - X (kN)

Peak Base 

Shear - Y (kN)

1 WSM80 468 467 274

2 SIN100(Y)-1 996 9 996

3 VOG75-1 869 682 665

4 VOG100 1003 831 754

5 VOG125 1163 979 870

6 VOG150 1317 1109 967

7 VOG175 1467 1237 1058

8 DIA80 1271 1064 916

9 DIA95(XY) 1457 1245 965

10 ELC130 851 677 719

11 IWA100(XY) 1212 766 1211

12 RRS88(XY) 1365 1100 1240

13 RRS88 1355 1097 1214

14 VOG75-2 808 645 624

15 SIN100(Y)-2 926 13 926

Table 5-3 Peak base shear coefficient for all simulations: total, x and y directions.

GM # GM Name Peak Base 

Shear Coeff.

Peak Base 

Shear Coeff. - X 

Peak Base 

Shear Coeff. - Y 

1 WSM80 0.09 0.09 0.05

2 SIN100 0.19 0.002 0.19

3 VOG75 0.17 0.13 0.13

4 VOG100 0.19 0.16 0.14

5 VOG125 0.22 0.19 0.17

6 VOG150 0.25 0.21 0.18

7 VOG175 0.28 0.24 0.20

8 DIA80 0.24 0.20 0.17

9 DIA95_2D 0.28 0.24 0.18

10 ELC130 0.16 0.13 0.14

11 IWA100 0.23 0.15 0.23

12 RRS88_2D 0.26 0.21 0.24

13 RRS88 0.26 0.21 0.23

14 VOG75 0.15 0.12 0.12

15 SIN100 0.18 0.002 0.18
77

Figure 5-6 Base shear coefficient for all simulations: total, x and y-directions

In Figure 5-7, the base shear coefficient for each motion is superimposed over the backbone 

force-displacement relation of the LR bearings, using the assumed design properties in Table 3-

3. The total force in the LR bearings was observed to exceed the design backbone for 

displacements less than 30 cm (12 in) and fall below the design backbone for displacements 

exceeding 30 cm (12 in). As an example, the base shear was approximately proportional to 

displacement as the intensity was increased from VOG75 to VOG175 simulations (simulations 3 

to 7 in Figure 5-7), but with a slope slightly lower than the second slopepost-yield stiffness Kd. 

The influence of ground motion intensity on the isolator response and modeling assumptions is 

addressed in Section 7.4.2. 

Most simulations followed this trend with the exception of the Rinaldi motions (simulations 12 

and 13, RRRS88(XY) and RRS88, in Figure 5-7). Assuming the design backbone curve was an 

accurate reflection of the bearing response, the observed points should fall below the design 

curve since the base shear coefficient represented an average bearing shear, while the 

displacement represented a peak displacement recorded in any LR bearing. An explanation for 

the higher observed base shear in the Rinaldi motions could not be found. 

0

0.05

0.1

0.15

0.2

0.25

Base Shear Coefficient

WSM80

SIN100(Y)-1

VOG75-1

VOG100VOG125VOG150VOG175

DIA80

DIA95(XY)

ELC130

IWA100(XY)

RRS88(XY)

RRS88

VOG75-2

SIN100(Y)-2

X Y Total
78

Figure 5-7 Base shear coefficient for each simulation alongside the design backbone 

curve

5.4 Axial Forces in LR Bearings

The static forces on the LR bearings at the start of the experiments were measured as: East = 

435 kN (98 kips), South = 755 kN (170 kips), North = 490 kN (110 kips), and West = 235 kN (53 

kips). As discussed in Section 4.3, the measured static loads on the bearings at the beginning of 

the experiments differed from the expected loads according to tributary area calculations. The 

actual measured and expected static loads on the LR bearings were compared in Table 4-2. In 

summary, the portion of the building weight carried by the LR bearings (about 37%) was 

significantly less than portion of the weight that was expected to be carried by the LR bearings 

(about 51%). The sources of the discrepancy could not be identified with certainty, but we 

mention the following probable causes: 1) The base of the testbed building was warped due to 

weathering and storage conditions. This caused the weight of the building to be distributed in a 

different pattern than if the building had been erected on top of the isolation system. 2) The 

stiffer CL bearings attracted more weight, thus carrying a larger portion than if the weight was 

balanced on a single type of isolator. Both factors were thought to contribute to the static load 

distribution measured at the part of the experiment. 

The axial forces in the LR bearings varied for each bearing and during each simulation due to a 

combination of factors including: variation in static forces, overturning, vertical excitation, and 

load transfer between LR bearings and CL bearings (discussed in Section 6.2). The peak 

compressive and tensile forces measured in any LR bearing for each simulation are shown in 

Figures 5-8 and 5-9, respectively. Tension was observed in at least one bearing for seven of the 

fifteen simulations (Figure 5-8). The largest compressive force in a single bearing was about 

2000 kN or 450 kips (about 40% of the static weight of the building) and the largest tensile force 

was 453 kN (102 kips), both observed during RRS88. The variation in axial force during RRS88 

was caused by the vertical excitation. In general, cavitation, or tensile rupture of the rubber 

0 10 20 30 40 50 60

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

1

2

3

4 5

6

7

8

9

10

11

12 13

14

15

Peak Displacement (cm)

Base Shear Coeff.

Design backbone curve

Test
79

matrix, is expected at a negative pressure = 3G (Constantinou et al. 2007), where G is the 

shear modulus of the rubber. Taking G to equal the design value of 0.41 MPa (0.06 ksi) (Table 

3-3), the approximate tensile force for cavitation in these LR bearings is 476 kN (107 kips), 

which only slightly exceeds the peak tensile force observed. Thus, the East LR bearing may 

have been on the verge of cavitation, or cavitation may have actually occurred, preventing the 

peak tensile force from going beyond the observed value.

Figure 5-10 illustrates the peak axial force in compression and tension for each LR bearing in 

each simulation, both absolute and normalized by the static force in the bearing at the start of 

the test program. Throughout the simulations, the South bearing was generally subjected to the 

largest compressive force, and the West bearing was subjected to the smallest compressive 

force (Figure 5-10(a)), which was in proportion to the static weight carried on the bearings. 

However, the normalized compressive force was largest in the West bearing, which carried the 

smallest static force, and smallest in the South bearing, which carried the largest static force 

(Figure 5-10(b)). Thus, the variation in compressive force, computed as a percentage of the 

static load, increased as the static load decreased. The largest tensile force generally occurred 

in the East LR (Figure 5-10(c)), which did not carry the greatest or least static force, but was 

usually subjected to the largest displacement (Figure 5-2). At large lateral displacements, a 

portion of the axial forces in the LR bearings were observed to transfer to the CL bearings, in 

some cases causing the LR bearings to be subjected to tension. Since the displacement 

demands were largest in the East bearing, the largest tensile forces occurred in the East 

bearing. The phenomenon of load transfer between LR and CL bearings is documented in 

Section 6.2.
80

Figure 5-8 Peak compressive force in any LR bearing for each earthquake simulation

Figure 5-9 Peak tensile force in any LR bearing for each earthquake simulation. (A 

tensile force of zero indicates that tension was not observed)

-2500

-2000

-1500

-1000

-500

0

Peak Compressive Axial Force (kN)

WSM80

SIN100(Y)-1

VOG75-1

VOG100VOG125VOG150VOG175

DIA80

DIA95(XY)

ELC130

IWA100(XY)

RRS88(XY)

RRS88

VOG75-2

SIN100(Y)-2

-1022

-824

-1209

-1318

-1433

-1541 -1629

-1459

-1312

-1150 -1101

-1201

-2021

-1267

-899

0

100

200

300

400

500

Peak Tensile Axial Force (kN)

WSM80

SIN100(Y)-1

VOG75-1

VOG100VOG125VOG150VOG175

DIA80

DIA95(XY)

ELC130

IWA100(XY)

RRS88(XY)

RRS88

VOG75-2

SIN100(Y)-2

62

149

279

142

280

17

453
81

Figure 5-10 Peak axial forces in each LR bearing for each simulation: (a) Peak 

compression force, (b) peak normalized compression force, (c) peak tension 

force, and (d) normalized peak tension force

5.5 Isolation System Re-Centering

The displacement recorded at the location of each LR bearing at the end of every simulation – 

referred to as permanent displacement - is shown in Figure 5-11. Prior to the 8th simulation 

(DIA80), the peak permanent displacement at any isolator location was about 0.5 cm (0.2 in). A 

sudden increase in the permanent displacement was observed in the x and y directions at the 

East and North LRB locations, respectively, at simulation #8. This permanent displacement 

recorded in the sensors reflected is believed to be a combination of permanent deformation in 

the bearings and sliding of the steel connecting plate.

From the inspection pictures taken at the end of the 1st day of testing which directly followed 

DIA80 (Figure 5-12), the bottom steel plate of the East LRB slid about 1.1 cm (0.4 in). It cannot 

be determined whether the slippage occurred during trial 8 or trial 7. However, as later shown in 

Section 6.1, slippage of the bolts connecting the LRB bottom steel plate to the supporting steel 

hexagonal plate of the load cell occurred as early as the 5th trial (VOG125), which led to the 

sliding of the bottom plate seen in Figure 5-12. If the sliding of the steel plate had not occurred, 

we believe that the permanent displacement in the bearings would have been limited to that 

observed in the first few simulations - around 0.5 cm (0.2 in) - which is insignificant.

Prior relaxation tests performed on LR bearings (Constantinou et al. 2007) suggested that the 

characteristic strength of LR bearings drops markedly under static conditions. Specifically, a 

bearing was returned to zero displacement following a sequence of large velocity cyclic loading 

and an imposed permanent displacement. When returned to zero displacement, the 

characteristic dropped to about 1/3 of its starting value after 8 minutes and 1/4 of its starting 

Static1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

-2500

-2000

-1500

-1000

-500

0

(a) Peak Compression Force

Peak Force (kN)

Static1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

1

2

3

4

5

(b) Normalized Peak Compression

Peak Force (kN)

Static1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

0

100

200

300

400

(c) Peak Tension Force

Peak Force (kN)

Static1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

0

0.2

0.4

0.6

0.8

1

(d) Normalized Peak Tension

Peak Force (kN)

E S N W
82

value after 30 minutes. The drop in characteristic strength due to relaxation suggests that 

residual permanent displacements in the bearings would disappear over time. 

The relaxation effect was evaluated in the present test program by considering any changes

looking for reductions in permanent displacement from the end of one simulation to the start of 

the next, which is illustrated separately for each bearing in Figure 5-13. Recall that the average 

time between simulations was about 50 minutes. Figure 5-13 does not indicate consistent 

reductions any significant changes in permanent displacements in the sensors from the end of 

one simulation to the start of the next that would be are consistent with a relaxation effect. 

Relaxation The changes in permanent displacement may have been inconsistent (sometimes

increasing and sometimes decreasing) may not have been observed since the bearing 

displacements, as computed by the string pots, were not independent but rather constrained to 

move together through the assumed base diaphragm constraint. Nonetheless, the observed 

permanent displacements were not significant.

Figure 5-11 X-direction, y-direction, and total (vector sum) displacement recorded in 

each LR bearing at the end of every earthquake simulation

REFLECTS BASE 

PLATE SLIPPAGE

REFLECTS BASE 

PLATE SLIPPAGE

REFLECTS BASE 

PLATE SLIPPAGE
83

Figure 5-12 Permanent displacement of around 1.1 cm on the East bearing due to 

sliding of the bottom steel plate
84

Figure 5-13 Permanent displacement at the location of each LR bearing at the end of 

one simulation compared to the beginning of the next

5.6 Floor Accelerations in the Testbed Building

The peak acceleration profile of the building (peak acceleration versus floor level) in both 

horizontal directions for all earthquake simulations is shown in Figure 5-14. Accelerations from 

multiple sensors were averaged as described in Section 4.6.3 based on the acceleration sensor 

layout in Figure 4-11. Although most individual simulations are not identified by input excitation, 

this plot format depicts the range of accelerations observed. The acceleration profile shape was 

similar for most excitations, which was almost linear from the base through the 4th floor followed 

by an increase in acceleration at the 5th and roof floors. The isolation system was very effective 

East Bearing

South

North

West
85

in attenuating the acceleration in the superstructure. Outliers are identified in Figure 5-14, which 

include on the low side: the service level motion Westmorland (WSM80) in both horizontal 

directions, and the sine wave simulations (SIN100(Y)- 1 & SIN100(Y)-2)) in the x-direction due 

to the unidirectional input; on the high side: Rinaldi (RRS88) in the y-direction as a result of the 

strong vertical input. The influence of vertical excitation on the overall response of the structure, 

with special attention to the RRS88 simulation, is explored in Chapter 10.

With the exception of RRS88, larger peak ground accelerations led to greater attenuation of 

acceleration as expected. Although it cannot represent variability due to ground motion 

frequency content, the reduction in floor accelerations relative to PGA is often used to quantify 

the effectiveness of the isolation system. During Diablo 95%, one of the largest motions applied 

to the system with PGA in the x-direction = 1g, the observed peak roof acceleration was 0.45g, 

which was a 65% decrease relative to PGA. Excluding the outliers, PGA ranged from around 

0.3g to 1g in the x-direction and 0.24g to 1.18g in the y-direction, while base level peak 

acceleration (just above the isolators) ranged from 0.19g to 0.5g in both directions. For Rinaldi 

88, the roof acceleration in the y-direction was greater than the PGA. 

To directly investigate the floor acceleration as a function of ground motion intensity, the floor 

acceleration profiles, both absolute and normalized by PGA, are plotted for the increasing 

intensity Vogtle excitations (VOG75, VOG100, VOG125, VOG150 and VOG175) in Figure 5-15. 

Recall that a low pass filter with a cutoff frequency of 50 Hz was applied to all signals (Figure 4-

5). The absolute floor accelerations increased consistently with increasing ground intensity, but 

the normalized accelerations decreased with increasing ground intensity as expected. 

The influence of vertical excitation is considered by comparing the absolute and normalized 

acceleration profiles for Diablo 95% (XY), Diablo 80% and Rinaldi 88% (XY and 3D) (Figure 5-

16). Even though the intensity of the Diablo 95% motion was substantially greater than the 

Diablo 80% motion, the floor accelerations were greater in Diablo 80%, which indicates that 

vertical excitation affected the recorded horizontal floor acceleration. For the Rinaldi motion, a 

significant amplification of horizontal floor acceleration was observed for 3D shaking relative to 

XY (horizontal only) shaking, which suggests a horizontal-vertical coupling phenomenon (see 

Chapter 10).
86

Figure 5-14 Peak acceleration profile for all simulations in both horizontal directions

 

Figure 5-15 Peak floor acceleration profiles for increasing intensity of Vogtle input 

excitation (75% - 175%)

0 0.5 1

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - X

Peak Acceleration (g)

0 0.5 1

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - Y

Peak Acceleration (g)

3D 2D & 1D

RRS88

RRS88

WSM80 WSM80

SIN100(Y)-1

and

SIN100(Y)-2
87

Figure 5-16 Peak absolute and normalized acceleration profile comparing XY and 3D 

excitations for Diablo (95% and 80% respectively) and Rinaldi (88%)

5.7 Story Drifts in Testbed Building

The peak story drift profiles (peak drift versus story level) in both horizontal directions for all 

earthquake simulations are shown in Figure 5-17. The drifts were calculated at the geometric 

center of each story level as described in Section 4.6.3. The peak drift in either direction 

occurred in the 2nd floor for all simulations, with the exception of both sine wave simulations, 

which incurred negligible drifts in the x-direction. The story drift decreased from the 2nd floor to 

the roof level, where the peak roof drift was generally less than the first story drift. 

The drifts for both RRS88(XY) and RRS88 in the y-direction were noticeably larger than the 

drifts observed in any other motion. This increase in drift was due to the predominance of low 

frequency components associated with the near-fault motion. The input acceleration history for 

the RRS88 simulation, shown in Figure 4-18, contains a strong pulse with a duration of about 1 

sec at the instant of peak acceleration in the y-direction. The relative intensity of ground 

acceleration in the x and y-directions was consistent with the trend of the drift profiles. Like the 

accelerations, the drifts increased significantly from RRS2D to RRS3D. The drift in the 5th floor 

was larger than the 4th floor in the y-direction for RRS3D, which is consistent with the 

acceleration profile of Figure 5-16.

0 0.5 1

Ground

Base

1st

2nd

3rd

4th

Roof

Peak Acceleration (g)

Acc. Profile - Diablo

0 0.5 1

Ground

Base

1st

2nd

3rd

4th

Roof

Acc. Profile - Rinaldi

Peak Acceleration (g)

0 0.5 1

Ground

Base

1st

2nd

3rd

4th

Roof

Normalized Acc. Profile - Diablo

Normalized Peak Acceleration

0 0.5 1

Ground

Base

1st

2nd

3rd

4th

Roof

Normalized Acc. Profile - Rinaldi

Normalized Peak Acceleration (g)

3D-X

3D-Y

2D-X

2D-Y
88

Figure 5-17 Peak story drift throughout the height of the building for all excitations in x 

and y directions

0 0.2 0.4 0.6

1st

2nd

3rd

4th

5th

Drift (%)

Story Level

Peak Story Drift - X

0 0.2 0.4 0.6

Peak Story Drift - Y

Drift (%)

3D 2D & 1D

RRS88

RRS88(XY)

WSM80

SIN100(Y)-1

SIN100(Y)-2

and
89

6. TECHNICAL RESPONSE OF HYBRID LR ISOLATION SYSTEM 

In this chapter, specific technical aspects germane to the response of the hybrid LR isolation 

system are explored. This is the first time that LR bearings and CL were tested as a hybrid 

isolation system on an earthquake simulator. Thus, unique load transfer between the two types 

of devices and stability aspects of the system are explored in depth. The topics discussed in this 

chapter include bolt slip, load transfer, and repeatability of the isolation system response after 

many tests. 

6.1 Bolt Slip in LR Bearings

Due to a variety of conditions unique to this experimental program, the bolted connections 

securing the LR bearings to the structure above and steel connecting plates below did not 

satisfy slip critical criteria, and slippage of the bolts was observed. The connections were 

anticipated to reach the slip critical limit at bearing displacements of about 40 to 45 cm (16-18 

in), but the bearings were tested out to displacements of 55 cm (22 in). In practice, the bearings 

would never be designed with low capacity at the connection level. The following factors 

influenced the connection design: 1) technical difficulties and prohibitive cost associated with 

drilling and tapping holes in the base of the testbed structure from beneath prompted the project 

team to select the smallest possible bolt size for the connection. 2) The bolt holes were 

oversized by 9 mm (0.4 in), deviating from standard practice, to accommodate ease of 

installation when lowering the testbed structure by crane over the 9 pre-installed isolators. 3) 

During pre-test planning and negotiation, when the connection design was finalized, it was 

doubtful that the bearings would be tested beyond 40 cm (16 in). Since bolt slip can easily be 

avoided in practice, its occurrence and subsequent influence on the response of the isolated 

building are documented briefly here for completeness.

As mentioned above, slippage was observed in the bolts that secured the LR bearing top 

connecting plate to the base of the structure above and the bottom connecting plate to the steel 

hex plate of the load cell assembly. Evidence of bolt slip included: 1) loud banging noise heard 

in-phase with the displacement cycles and subsequently observed instantaneous force drops 

and/or spikes in the LR bearing forces recorded by the load cells, and 2) movement of the LR 

bearings relative to the structure above and below observed in post-test inspection, which was 

shown in Figure 5-12 and is further illustrated in Figure 6-1 below. 

The bolt slip was first observed during Vogtle 125%, and continued to be observed throughout 

the simulation sequence, wherein larger displacements in the bearings increased the instances 

of bolt slip. Some slip was observed in all four LRBs, but the largest force spikes and drops 

occurred in the East bearing. Figures 6-2 plot snapshots of the unfiltered force history of the 

East bearing (LRB-E) in the x and y-directions for 8 of the 15 simulations, which are labeled by 

trial number, the 3 letter abbreviation for the ground motion, and the scale factor. (Recall that, 

as stated in Chapter 4, all data shown is filtered unless otherwise indicated.) In these figures, 

the force drops are first observed in Vogtle 125% around 11.5 sec and 12 sec. In the next trial 

(Vogtle 150%), additional drops are observed at other time instances, and spikes are observed 

at 11.5 and 12 seconds. Then, in Vogtle 175%, even more spikes and drops are observed. After 

Vogtle 175%, force drops and spikes continue to appear (e.g. Diablo 80% at about 15 sec) but 

with decreasing intensity. The drops and spikes are also observed in the bearing hysteresis 

loops, such as those plotted for LRB-E during Vogtle 125% and Vogtle 150% (Figure 6-3). The 

force drops and spikes tend to be observed during large displacement cycles just before the 

peak displacement is reached.
90

Figure 6-1 Movement of (a) East LR bearing relative to (b) top and (c) bottom plates.

Top plate 

slippage

I

II

Base plate 

slippage

(a)

(b) (c)

I II
91

Figure 6-2 Horizontal force history of the East bearing (LRB-E) in the x- and ydirections for a subset of the trials

The horizontal and vertical force histories for all four LR bearings are shown in Figure 6-4 for 

Vogtle 150%, which demonstrates that the greatest amount of bolt slip occurred in LRB-E. 

During Vogtle 150%, LRB-E is the only bearing that experienced both substantial force spikes 

and force drops. Several force drops are observed in LRB-W, but they are small in intensity 

compared to LRB-E. Only one small drop/spike for LRB-N (around 12 sec), and two in LRB-S 

are evident (Figure 6-4).

The synchronized vertical force histories of each bearing are also plotted in Figure 6-4 to 

provide additional insight as to why the bolt slip may have occurred. One proposed theory is that 

bearing tension contributed to the bolt slip. The addition of CL bearings to the isolation system 

did not entirely prevent tension in the LR bearings, which is discussed in the next section. 

During the Vogtle 150% record, tension (bearing vertical force greater than zero in Figure 6-4) is 

observed more frequently and with larger intensity in LRB-E and LRB-W, which also have the 

-500

0

500

VOG75-1 VOG175

-500

0

500

VOG100 DIA80

-500

0

500

VOG125

Isolator Force (kN)

DIA95(XY)

10 15 20

-500

0

500

VOG150

Time (s)

10 15 20

VOG75-2

Time (s)

x

y
92

most obvious horizontal force drops/spikes. The instances of bolt slip seen in this figure do not 

align with the instances of peak tension, but horizontal force drops/spikes always occur after 

tension has been observed in the bearing. At the same time, LRB-E is subjected to the largest 

displacement demands (42.3 cm (16.7 in) in Vogtle 150% compared to 35.7 cm (14.1 in) in 

LRB-W), and thus experiences a larger shear force that makes it more susceptible to bolt slip.

Figure 6-5 plots the unfiltered force in LRB-E and unfiltered accelerations in the SE column 

sensors at all floors in x and y-directions, respectively, for Vogtle 150%. In this figure, a spike in 

the bearing horizontal force is always preceded by a small force drop. The following explanation 

is consistent with the drop/spike pattern. After a large tension excursion, the bolts start to slip 

and move with respect to the oversized holes. The bolt movement stops the bearing movement, 

causing it to instantaneously unload, corresponding to the first drop in force. If the bolts reach 

the other side of the oversized hole, an impact occurs, resulting in a force spike and a loud 

banging noise. Due to the instantaneous nature of the impact, dynamic amplification occurs, 

causing what appears to be an instantaneous spike/drop, but is actually very high frequency 

oscillation. 

Figure 6-5 also shows that the bolt slip induces a dynamic amplification in the floor 

accelerations that diminishes with increasing height in the building. The acceleration spikes 

appear to be timed with the first force drop associated with the start of bolt slip and not the 

second force spike/drop associated with impact of the bolt against the other side of the hole. In 

the overall test program, bolt slip (by itself) did not appear to affect the performance of 

nonstructural components or cause disruption of contents, located on the 4th floor and above. 

There is no evidence of whether nonstructural components on the lower floors would have been 

affected by the bolt slip. Filtering the recorded force and acceleration data significantly reduced 

the drops and spikes resulting from bolt slip, but did not completely eliminate them.

In summary, in this experimental program, the bolt slip did not appear to affect the response of 

the isolation system aside from the drops/spikes in force, and the adverse effects on the 

structural response were limited. However, the possibility that increased acceleration would 

affect the response of nonstructural components and contents or compromise performance in 

any way is an unnecessary risk. The observations from these experiments reinforce the 

conclusion that bearings should always be designed with slip critical connections, as they 

routinely are in practice. 
93

Figure 6-3 X and y-direction hysteresis loops (horizontal force vs. displacement) of the 

East bearing (LRB-E) during Vogtle 125% and Vogtle 150%

-800

-600

-400

-200

0

200

400

600

800

VOG125 - X

Isolator Force (kN)

VOG150 - X

-40 -20 0 20 40

-800

-600

-400

-200

0

200

400

600

800

VOG125 - Y

Displacement (cm)

Isolator Force (kN)

-40 -20 0 20 40

VOG150 - Y

Displacement (cm)
94

Figure 6-4 Horizontal (Fx and Fy) and vertical (Fz) force history of all four LR beaings

during Vogtle 150%

-1600

-1000

-600

-300

0

300

-300

0

300

-300

0

300

600

900

LRB-E

Isolator Force (kN)

-1600

-1000

-600

-300

0

300

-300

0

300

-300

0

300

600

900

LRB-S

10 15 20

-1600

-1000

-600

-300

0

300

-300

0

300

-300

0

300

600

900

LRB-N

Time (s)

Isolator Force(kN)

10 15 20

-1600

-1000

-600

-300

0

300

-300

0

300

-300

0

300

600

900

LRB-W

Time (s)

Fx

Fy

Fz
95

Figure 6-5 Propagation of bolt slip through the height of the structure in x-direction 

during Vogtle 150%; unfiltered horizontal force in LRB-E and 1st – 6th floor 

acceleration in SE column

9 10 11 12 13 14 15 16

-50

0

50

-1

0

1

-1

0

1

-1

0

1

-1

0

1

-1

0

1

-1

0

1

Time (s)

Force (kN)

1st fl acc (g)

2nd fl acc (g)

3rd fl acc (g)

4th fl acc (g)

5th fl acc (g)

6th fl acc (g)
96

6.2 Transfer of Load Between Bearings

Axial force transfer between the LR bearings and the CL bearings over the course of the 

simulations was expected. LR bearings, when subjected to combined compressive load and 

lateral displacement, reduce in height, as shown in Figure 6-6(a), where P is the axial force and 

δ is the downward deflection or shortening. However, downward movement of the LR bearings 

is constrained by the rigidity of the base diaphragm and the axial stiffness of the CL bearings, 

which is about 2.5 times the stiffness of the LR bearings. The constraint generates an upward 

force F on the bearings (Figure 6-6(b)), which This causes load to redistribute from the LR 

bearings to other isolators. If P exceeds F, The net effect is a reduction in axial compressive 

force carried by the LR bearings occurs, while if F exceeds P and individual the LR bearings will

may go into tension to satisfy the base diaphragm constraint. This type of behavior can occur at 

large displacements and is (unrelated to system overturning) to satisfy the base diaphragm 

constraint. We refer to this subsequently as the “load transfer” effect. 

Figure 6-6 Resultant action on LR bearings as a result of CL bearings and base 

diaphragm constraint

Evidence of the load transfer effect was observed during the test program. Histories of isolator 

displacements and axial forces on individual LR bearings and summed over all LR bearings are 

shown for three different XY excitations: sine wave input (Figure 6-7), Diablo 95% (Figure 6-8) 

and Rinaldi 88% (Figure 6-9). The displacements shown have been computed by averaging 

displacements of LRB-E and LRB-W in the x-direction, and LRB-N and LRB-S in the y-direction; 

axial force is considered to be positive in tension. Recall that forces acting on the CL bearings 

were not measured during the experimental program. The sine wave simulation is the simplest 

to interpret because the input to the building was unidirectional in the y-direction, generating 

very little torsional response in the isolators. Vertical lines drawn through local (vector sum) 

peak displacements and extended through the axial force plots demonstrate that every time a 

peak displacement is reached (either local maximum or local minimum), a corresponding net 

reduction in total axial force of the 4 LR bearings (black line in Figure 6-7) is observed. The axial 

forces in individual LR bearings are more complex since overturning effects are present. LRB-N 

and LRB-S, which being close to the neutral axis of the building plan for y-direction input should 

not experience much overturning, also appear to consistently unload at every local 

displacement peak – max or min (Figure 6-7). The instant of least compressive axial force in the 

bearings does not exactly correspond to the instant of peak displacement shown, but is close 

enough that the load transfer trend is confirmed. With regard to individual bearings, LRB-W 

sustains maximum compression for displacements in the positive y-direction and minimum 

compression for displacements in the negative y-direction, while LRB-E experiences the 

opposite, which is the expected trend when subjected to overturning related axial force 

demands (Figure 6-7). Thus, for this excitation, the overturning effect in LRB-E and LRB-W is 

P δ

F-P

(a) with shortening (δ) (b) without shortening
97

stronger than the load transfer effect. However, fluctuation of the axial force between the 

displacement peaks suggests that both the overturning effect and the load transfer effect are 

contributing to the response.

Figure 6-7 History of average horizontal displacement (x, y and vector sum), and axial 

force in individual LR bearings and summed over all LR bearings for Sine 

Wave (XY)

For Diablo 95% (XY) (Figure 6-8), substantial load transfer, as indicated by axial unloading of 

the LR bearings, occurs at 4 different time instants corresponding to peak displacements 

(combination of x and y-direction movement) observed at the center of the building. LRB-E 

sustains tension at every one of these time instants, and at two different instants (just after 15 

sec and about 19 sec), the total axial force on the LR bearings exceeds 0, indicating that the 

entire weight of the building has shifted to the CL bearings. The load transfer effect is much 

0

30

-30

0

30

-30

0

30

Disp (cm)

Sine 100% (Y) - 1

x-dir

y-dir

vector sum

5 6 7 8 9 10 11 12 13 14 15

-2500

-2000

-1500

-1000

-500

0

500

Axial Force (kN)

Time (sec)

S E W N Total
98

more significant for this simulation than the sine wave since the isolator displacement is much 

larger (55 cm or 22 in compared to 21 cm or 8 in, see Figure 5-1). As a result of the torsional 

demand on the isolation system discussed previously, the largest displacements are 

consistently observed in LRB-E, which is also subjected to the greatest tension. 

Figure 6-8 History of average horizontal displacement (x, y and vector sum), and axial 

force in individual LR bearings and summed over all LR bearings for Diablo 

95% (XY)
99

Figure 6-9 History of average horizontal displacement (x, y and vector sum), and axial 

force in individual LR bearings and summed over all LR bearings for Rinaldi 

88% (XY)

To understand the extent of load transfer when the isolator displacement is 55 cm, consider 

than the overlapping area Ar between the top and bottom areas of a circular bearing at a given 

displacement is calculated by (AASHTO, 2010):

δδ = 2 cos−1 �

dd mmmmmm

DD � = 2 ∗ cos−1 � 55cm

69.85cm� = 1.34 (6.1)

AArr = DD2

4 (δδ − sinδδ) = 69.8524 (1.34 − sin(1.34)) = 449.2cm2 (6.2)

where D is the bearing diameter and dDmax is the peak displacement. At the peak recorded 

displacement of 55 cm (22 in), the overlapping area is a small fraction – about 12% – of the total 

bonded area (3832 cm2 or 592 in2

). According to the overlapping area rule (Buckle and Liu 
100

1994), the bearing has sustained an 88% loss in axial force capacity, which confirms that 

upward forces are generated to counteract the natural shortening in the bearing. The 

displacement pattern shown in the time series plots of Figure 6-8 (also in Figure 5-3) indicates 

that the peak displacement cycle occurs along a diagonal (from NE corner to SW corner). Since 

the load transfer is partially counteracted by the effects of overturning, LRB-N and LRB-E 

sustain less load transfer (axial unloading) for a positive excursion in x and a negative excursion 

in y, and LRB-S sustains less load transfer for a negative excursion in x and a positive 

excursion in y (Figure 6-8). The trend for LRB-W is inconclusive. The load transfer effect may be 

stronger than the overturning effect in LRB-W because it carries significantly less static weight 

than the rest of the LR bearings (Section 5.4). The load transfer effect is also observed during 

Rinaldi 88% (Figure 6-9), although the less axial force unloading occurs due to the lower peak 

displacement in Rinaldi (38 cm or 15 in from Figure 5-1). 

While the transfer of load between LR bearings and CL bearings is evident for XY simulations, 

the trends are more difficult to ascertain in 3D simulations that include vertical excitation. In this 

series, Rinaldi 88% is the only excitation applied both with and without vertical input, thus 

allowing for the effects of vertical shaking to be directly assessed. Figure 6-10 shows the axial 

force histories for each of the four bearings and the total for Rinaldi 88%. By inspection, the 

axial force histories for 3D excitation are rich in higher frequency content compared to the force 

histories for XY excitation. Thus, instances of load transfer are less obvious. However, from a 

direct comparison of the axial forces in Rinaldi 88% (XY) and Rinaldi 88% (Figure 6-11), the 3D 

forces oscillate about the backbone of the XY forces. Thus, the 3D force variation is essentially 

equal to the XY force variation augmented by an additional high frequency component. To 

eliminate the force variation due to vertical excitation, and thus verify the pattern of load transfer 

for the 3D simulation, a low-pass Butterworth filter with a cutoff frequency of 2 Hz is applied to 

the total axial force for Rinaldi 88%. This filter has the same shape as that shown in Figure 4-5 

when normalized with respect to the cutoff frequency. The 2 Hz cutoff frequency was selected 

since it preserves the frequencies related to horizontal vibration of the isolation system but 

eliminates typical frequency of vertical excitation and response. The resulting filtered axial force 

is shown in Figure 6-11 as a red dashed line superimposed over the unfiltered total axial force. 

The filtered total axial force for 3D excitation matches that for XY excitation very closely.

Knowing that the trend for 3D excitation can be identified, an XY versus 3D comparison is 

attempted for the Diablo excitation, where the 3D simulation data is available at a different scale 

factor (80%) than the XY simulation (95%). The total (vector sum) displacement (computed as 

described for Figures 6-7 to 6-9, axial forces in each LR bearing, and total axial force (with and 

without filtering for the 3D simulation) are compared for Diablo 95% (XY) and Diablo 80% in 

Figure 6-12. The peak displacement demand at the center of the building (Figure 6-12) is about 

25% lower in Diablo 80% than Diablo 95% (XY). However, in Diablo 80%, about half of the load 

transfers from the LR bearings to the CL bearings, while in Diablo 95% (XY), all of the load 

transfers from the LR bearings to the CL bearings at two different time instances. This indicates 

a nonlinear or escalating trend in the amount of load transfer with increasing horizontal 

displacement. The bearing tensile force demands would need to be analyzed prior to executing 

Diablo 95% as a 3D simulation; as it stands, almost no tension was observed in Diablo 80% 

(Figure 6-12).

In the context of the previous information, horizontal displacement and axial force data is 

presented for Vogtle 175% (Figure 6-13), which represents the largest simultaneous horizontal 

displacement and vertical excitation. The pattern of load transfer for individual LR bearings is 

evident even without filtering, and applying filtering to the total force confirms the pattern. Peak 

tensile demands in individual bearings are not as great for Vogtle 175% as they were for Diablo 

95% (XY), for which the bearing displacements are largest, or Rinaldi 88%, for which the vertical 

excitation input is largest.
101

Figure 6-10 History of axial force in individual LR bearings and summed over all LR 

bearings for Rinaldi 88%
102

Figure 6-11 History of axial force in individual LR bearings and summed over all LR 

bearings comparing Rinaldi 88% – dashed line – and Rinaldi 88% (XY) – 

solid line. A low pass filter with a cutoff frequency of 2 Hz was applied to the 

total axial force for Rinaldi 88% and is superimposed over the total, shown 

as a red dashed line

-800

-400

0

400

-800

-400

0

400

-800

-400

0

400

-800

-400

0

400

Axial Force (kN)

Rinaldi 88% (XY) vs Rinaldi 88%

S E W N

5 10 15 20

-8000

-6000

-4000

-2000

0

2000

Axial Force (kN)

Time (sec)

total 3D total 2D 3D w/filter
103

Figure 6-12 History of average vector sum displacement at building center, axial force in 

individual LR bearings and axial force summed over all LR bearings 

comparing Diablo 95% – dashed line – and Diablo 80% (XY) – solid line. A 

low pass filter with a cutoff frequency of 2 Hz was applied to the total axial 

force for Diablo 80% and is superimposed over the total, shown as a red 

dashed line

0

50

Diablo 95% (XY) vs. Diablo 80%

Vector Disp (cm)

3D XY

-800

-400

0

400

-800

-400

0

400

-800

-400

0

400

-800

-400

0

400

Axial Force (kN)

S E W N

10 15 20 25

-4000

-3000

-2000

-1000

0

Axial Force (kN)

Time (sec)

total 3D total 2D 3D w/filter
104

Figure 6-13 History of average vector sum displacement at building center, axial force in 

individual LR bearings and axial force summed over all LR bearings for 

Vogtle 175%. A low pass filter with a cutoff frequency of 2 Hz was applied to 

the total axial force and is superimposed over the total, shown as a red 

dashed line

While we have demonstrated that force transfers between LR bearings and CL bearings over 

the course of a simulation, a different pattern emerged between simulations, which is illustrated 

in Figure 6-14. This figure shows that the total compressive force carried by the LR bearings 

always increases from the beginning to the end of the simulation, which means that a small 

amount of load transfers from the CL bearings to the LR bearings over the course of the 

simulation. However, from the end of one simulation to the beginning of the next, the total 

compressive force carried by the LR bearings is consistently observed to decrease (Figure 6-

14(a)), which suggests that the original compressive force on the LR bearings would be restored 
105

over time. As mentioned before, the time between simulations in a given day was about 50 

minutes. At the end of Day 1, the total compressive force increase on the four LR bearings is 

about 430 kN or 97 kips (about 22% of the original total of 1915 kN or 431 kips). As confirmation 

of the relaxation effect, the original static load on the LRBs is nearly restored by the start of Day 

2; only about 50 kN (11 kips) additional remained (Figure 6-14(a)). Nearly the same pattern is 

repeated on Day 2, except that additional compressive force accumulates more quickly as 

larger motions were executed earlier in the day. 

Our hypothesis regarding the pattern of axial force transfer between simulations is as follows. 

When the bearings are constrained at large lateral displacements as depicted in Figure 6-6, the 

lead is sheared laterally and takes the shape of the slanted cylinder. The height of the slanted 

cylinder is longer than the height of the cylinder in the undeformed configuration. Upon return to 

the original position, the slanted cylinder tries to rotate, but the axial load and confinement from 

the shims force the lead cylinder back to its original height. Immediately upon return to the 

undeformed configuration, the lead plug exerts some force onto the upper and lower plates, or 

alternatively a slight height increase is maintained if the upper and lower plates are not 

confined. The confining force or height increase causes the LR bearings to temporarily take on 

additional load. After a short period of time, the force relaxes and the original state is restored. 

This effect would not be seen if a) the hybrid isolation system used elastomeric bearings instead 

of LR bearings, b) the base diaphragm was less stiff, or c) the system consisted entirely of LR 

bearings.

The maximum increase in the static compressive force in the LR bearings due to the effect 

described above is about 500 kN, which is about a 25% increase relative to the static load 

carried by the LR bearings at the start of the test program. Figure 6-14(b) provides a visual 

impression of the total static force variation throughout the test program.

As described in Section 3.2.2, the combination of LR and CL bearings was necessary in this 

program to provide a sufficient period shift and displacement capacity for the relatively 

lightweight structure. Chapters 8-10 describe numerical simulation to validate and demonstrate 

the predictability of the experimental response. However, we do not attempt to predict the axial 

forces in the bearings as part of this simulation. The forces in CL bearings cannot be validated 

since they were not measured. 

The observed load transfer in these experiments suggests that the compliance of the devices in 

a hybrid system must be carefully considered. The experiments have demonstrated that when 

working with a high stiffness tension capable device, the LR bearings can be subjected to nonnegligible tension due to a combination of load transfer and vertical excitation; the tension is not 

related to overturning. While the load transfer and the resulting tension observed in the LR 

bearings in these experiments is considered acceptable, the main drawback is that the peak 

force demands in the bearings (both tensile and compressive and in both types of devices) are 

difficult to predict; see Chapter 8-10 for further discussion. This is a concern since reliable 

numerical simulation is a requirement for the design of a nuclear structure. 

Due to these concerns regarding the load transfer, it is pertinent to consider the likely 

differences in response between the tested hybrid system and an isolation system composed 

exclusively of LR bearings. First, suppose the CL bearings were removed from the test setup, 

and the as-designed LR bearings were located beneath the four corner columns. Further, 

suppose that the same DBE and beyond DBE motions were posed that produced displacement 

demands up to 30 cm and 55 cm, respectively. The likely differences in the response of the 

exclusive LR system and the hybrid LR are discussed as follows.
106 

Figure 6-14 (a) Relative change in compressive force and (b) Total compressive force 

summed over LR bearings at the start and end of all simulations

The effective isolation properties and displacement demands for a DBE are expected to be 

essentially unchanged for an exclusive LR system. The bearing hysteresis loops in the 

exclusive LR system and the hybrid LR system are expected to look very similar. 

Without CL bearings, individual LR bearings are estimated to sustain overturning induced 

tensile loads on the order of 200 kN (see Equation (3.4) and related discussion). Based on 

Figure 5-10, a tensile load on the order of 350 kN was observed in one bearing in this 

experiment due to the combined effects of load transfer and vertical excitation. Vertical 

excitation would also increase the peak tensile load in an exclusive LR system, such that the 

peak tensile demands on individual bearings in the hybrid LR system and in the exclusive LR 

system are expected to be similar.

End of day 1

(a)

(b)
107

If not constrained by the hybrid setup, a simple analysis predicts that one or more bearings 

may experience a complete loss of horizontal stiffness at the displacement demands of beyond 

DBE motions (55 cm). As discussed earlier, the bearing overlapping area at 55 cm of 

displacement is about 12% of the total bonded area (Equation (6.2)), and thus by the 

overlapping area rule is predicted to reduce to 12% of the bearing critical buckling load in the 

undeformed configuration (Buckle and Liu 1994). The nominal critical buckling load of the

bearings is Pcro = 12,600 kN (Figures 3-8 and 3-9 for LR Option 3), and thus the reduced 

buckling load is Pcr = 1500 kN at a displacement of 55 cm. With only 4 LR bearings, the 

average static load of P = 1300 kN per bearing nearly exceeds the reduced critical buckling 

load, and thus the bearings are expected to buckle if subjected to compressive force increase 

due to overturning. In this case, the reduced horizontal stiffness of the bearing Kb of the bearing 

as a function of the nominal stiffness Kbo

2

1 cr

P

b bo P

K K        

  = −     (6.3)

would also tend to zero. As a worst case scenario, buckling and temporary loss of stiffness in 

one or more bearings could cause a global collapse of the system.

However, emerging studies suggest that the stability capacity of elastomeric bearings at large 

displacements is much larger than predicted by the overlapping area rule, and the isolation 

system has remarkable ability to recover from local instabilities in one or more bearings. For 

example, Sanchez et al. (2012) subjected a rigid block system with 4 one quarter scale natural 

rubber bearings to ground motions that imposed bearing displacements beyond their 

theoretical and experimentally observed stability limits. In one instance, one of the bearings 

was driven to a displacement 1.2 times its diameter and about twice the displacement at which 

loss of stiffness was observed, and the composite isolation system had a large negative 

stiffness. The isolation system successfully recovered from this and other excursions into the 

instability range. In an experimental study of a two-fifth scale 3-span horizontally curved girder 

bridge with two isolation bearings at each bent and abutment, isolators were shown to remain 

stable at displacements 1.33 times the bearing diameter (Monzon et al. 2013). At a slightly 

larger displacement (1.4 times the bearing diameter), both bearings at one abutment 

experienced a local instability. One bearing was observed to sit down on the bottom plate while 

the top plate touched the side of the bearing. The other isolators remained stable because the 

bent bearings were larger diameter and because the displacement demands at the abutment 

were larger due to asymmetry in the curved bridge. The system recovered from instability 

multiple times, and no change to the isolator hysteretic properties were observed following 

instability. 

Relating these results to a hypothetical test of an exclusive LR isolation system at E-Defense, 

another possible scenario is that the isolated structure could survive excursions to 

displacements equal to the bearing diameter (70 cm) and beyond without collapse due to the 

ability to recover from local instabilities. The potential for recovery would be aided by the fact 

that displacements on one of the building were consistently observed to be substantially 

smaller than displacements on the other side due to torsion, and that axial loads on one side of 

the building would be less than the static loads due to overturning. Both of these variations 

increase the likelihood that the instabilities are localized, enhancing the chance of a global 

recovery. As a result of localized instabilities, the bearing hysteresis loops would change 

substantially compared to the ones observed in these tests, and exhibit negative stiffness loops 

at large displacements, similar to those observed in Sanchez et al. (2012). 
108

Next, suppose instead that the isolation system were designed without the constraints of the 

test setup for an actual nuclear facility. Removal of the test constraints implies the following 

could be achieved. Individual bearings would carry greater static loads, and thus their size (both 

diameter and bearing height) could be increased without altering the design properties (period

and damping) of the system. As a result of the size increase, the target displacement demand 

for the Vogtle site (about 60 cm) could easily be accommodated without approaching the 

stability limit of the bearing. As a result, stable hysteresis loops similar to the ones observed in 

these tests (discussed further in Chapters 7 and 9) would be expected. Assuming the nuclear 

facility is squat in plan compared to the tested 5-story steel frame building, the isolation system 

could feasibly be designed to eliminate overturning-induced tensile demands. Incidental bearing 

tension due to vertical excitation would still be likely, and the influence of vertical excitation on 

the response of the system (see Chapter 10) would be unaffected. Because a safety related 

nuclear facility could be designed exclusively with LR bearings without the need for CL 

bearings, the constraints encountered in the test program would not be encountered in the 

design of a nuclear facility, thus the hybrid LR isolation system tested at E-Defense need not be 

used for safety related nuclear facilities. As such, the load transfer issues that were apparent in 

the tests are of no consequence to practical applications.

6.3 Repeatability of Isolation System Response after Many Tests

During the test program, repetitions were conducted for two of the input ground excitations: Sine 

100% (Y) and Vogtle 75%. Each of these excitations was applied near the beginning of the test 

program and then again at the end of the test program for the hybrid LR isolation system. The 

repetitions were planned to assess the consistency of the isolation system response after the 

LR bearings had been subjected to many ground motions and to very large ground motions.

The variation in the input signals for these repetitions is assessed in Figures 6-15 and 6-16. 

Figure 6-15 compares the x and y-direction feedback acceleration at the center of the 

earthquake simulator (averaged over several accelerometers) for the first and second Sine 

100% (Y) and Vogtle 75% simulations. This feedback acceleration is the input acceleration to 

the isolated structure. Figure 6-16 compares the comparable 5% damped response spectra for 

the feedback acceleration. Figures 6-15 and 6-16 demonstrate that the input accelerations for 

the first and second simulations are essentially identical for both Sine 100% (Y) and Vogtle 

75%.
109

Figure 6-15 Sample feedback acceleration at the center of the earthquake simulator 

(averaged over several accelerometers) in the x and y-directions compared 

for the first and second Sine 100% (Y) and Vogtle 75% simulations
110

Figure 6-16 5% damped response spectra for the feedback acceleration at the center of 

the earthquake simulator (averaged over several accelerometers) in the x, y 

and z-directions compared for the first and second Sine 100% (Y) and Vogtle 

75% simulations

The Sine 100% (Y) excitation is a unidirectional input consisting of several similar amplitude 

sinusoidal cycles used for characterization of the isolation system. As mentioned in the previous 

section, the input was applied in the y-direction to minimize torsional response. Minor 

displacements and forces in the x-direction were observed that had no distinguishable effects 

on the y-direction hysteresis loops. The y-direction bearing force versus displacement 

(hysteresis loops) and y-direction displacement/force histories of all LR bearings are compared 

for the two Sine 100% (Y) simulations in Figures 6-17 and 6-18, respectively. From the 

hysteresis loops, a small decrease in bearing force is observed in the second simulation relative 

to the first (Figure 6-17). This force decrease is most notable as the displacement peak is 

approached in the negative direction. The peak force cycle-by-cycle is also slightly lower for the 

second simulation than the first based on the force histories (Figure 6-18), which is true for all 

0

0.5

1

1.5

2

2.5

X-dir

Spectral Acceleration (g)

0

0.5

1

1.5

X-dir

0

0.5

1

1.5

2

2.5

Y-dir

Spectral Acceleration (g)

0

0.5

1

1.5

Y-dir

10-2 10-1 100 0

0.5

1

1.5

2

2.5

Z-dir

Spectral Acceleration (g)

Period (sec)

10-2 10-1 100 0

0.5

1

1.5

Z-dir

Period (sec)

SIN100(Y)-1 SIN100(Y)-2 VOG75-1 VOG75-2

10-2 10-1 100 0

0.5

1

1.5

Z-dir

Period (sec)

10-2 10-1 100 0

0.5

1

1.5

2

2.5

Z-dir

Spectral Acceleration (g)

Period (sec)

0

0.5

1

1.5

Y-dir

0

0.5

1

1.5

2

2.5

Y-dir

Spectral Acceleration (g)

0

0.5

1

1.5

X-dir

0

0.5

1

1.5

2

2.5

X-dir

Spectral Acceleration (g)
111

LRBs. The observed decrease in the peak force is on the order of 5 to 11% when all LR 

bearings are considered, which is small.

The following hypothesis is offered to explain the mild variation in bearing response over the two 

simulations. The characteristic strength of LR bearings is affected by heating and temperature 

increase in the lead plug (Kalpakidis and Constantinou 2009a). Specifically, the characteristic 

strength of the lead plug decreases as the temperature increases under repeated cycling. While 

repetitions performed with the same starting temperature should produce identical response in 

the bearings, repetitions performed at different starting temperatures would produce slightly 

different response in the bearings. Since the test program consisted of many simulations in a 

single day with only 45-50 minutes between simulations, it is likely that 1) the bearings did not 

have time to cool to the starting temperature between simulations, and 2) the starting 

temperature for each simulation was different based on the recent simulation history. Consistent 

with the observations, the starting temperature for the second SIN100 simulation was likely 

higher than for the first SIN100 simulation, since the second followed a full day of testing 

including DIA95(XY), which generated the largest displacement demand, and IWA100(XY), 

which was more than 4 minutes long.

Although the forces decrease, the overall displacement demands do not increase in the second 

Sine 100% (Y) simulation. Rather, the hysteresis loop for each bearing shifts slightly to the left 

in the second simulation so that the negative direction peak increases but the positive direction 

peak decreases. This shift can be observed in every bearing hysteresis loop except for LRB-S 

(Figure 6-17), and in the displacement histories (Figure 6-18). The shift could be related to a 

small residual (permanent) displacement present at the beginning of the second Sine 100% (Y) 

simulation in some bearings; residual displacement is visible only in LRB-S and LRB-N in the 

opposite directions (Figure 6-18), which indicates that the permanent displacement is torsional. 

Recall that a portion of the permanent displacement is believed to be a result of sliding in the 

connection plates associated with bolt slip (see Sections 5.5 and 6.1). The absolute peak 

displacement (observed in the negative direction) increases on the order of 3 to 5% in the 

second simulation when all LR bearings are considered, which is insignificant.
112

Figure 6-17 Y-direction force vs. displacement (hysteresis loops) for the 1st and 2nd

simulation of Sine 100%(Y)

Additional plots are included to evaluate the consistency of the bearing response in the Vogtle 

75% simulations, for responses in both directions. X and y-direction bearing force versus 

displacement (hysteresis loops) are compared for the Vogtle 75% simulations in Figures 6-19 

and 6-20, while x and y-direction displacement/force histories of all LR bearings are compared 

in Figure 6-21 and 6-22. An additional figure (Figure 6-23) indicates the displacement traces 

(displacement in x versus displacement in y) of all LR bearings for the two Vogtle 75% 

repetitions. This figure is necessary to evaluate the consistency of the peak displacement, which 

is a vector quantity for bidirectional horizontal excitation, over the two Vogtle simulations. 

Again, by visual inspection of the hysteresis loops, the forces in the LR bearings decrease 

slightly for the second repetition of Vogtle 75% relative to the first (Figures 6-19 and 6-20). This 

reduction in force seems to be smaller for Vogtle 75% than for Sine 100% (Y), but a statistical 

evaluation was not completed. The absolute peak displacement increases for each LR bearing 

in the second simulation relative to the first, but also appears to result from the entire hysteresis 

loop shifting to the direction of negative displacement rather than a true increase in the 

displacement demand. The increase in absolute peak displacement for the second simulation of 

Vogtle 75% relative to the first is about 8% in all LR bearings (determined from Figure 6-23). 

Force and displacement histories are very similar when superimposed over each other for the 

first and second simulations (Figures 6-21 and 6-22). 
113

Figure 6-18 Y-direction displacement and force histories for the 1st and 2nd simulation of 

Sine 100% (Y)

The peak force in the LR bearings also decreases from cycle to cycle while the displacement 

remains approximately constant over the course of a simulation. This behavior is also indicative 

of dynamic reduction in yield strength due to heating of the lead plug. This behavior is observed 

most clearly during the Sine 100% (Y) simulation, which contains several regular displacement 

cycles of similar amplitude. The hysteresis loops generated during Sine 100% (Y), both first and 

second simulations, are plotted again for LRB-N in Figure 6-24, where the cycles are 

individually identified. In this figure, the forces decrease in every cycle, with the greatest 

reduction between the first and the second cycle. While the total reduction in strength over the 8 

cycles of the sine wave is significant, a typical motion includes at most a few cycles of strong 

amplitude motion. Thus, a reduction in strength is not noticeable for the majority of the motions 

included in the test program, such as Vogtle 75% (Figures. 6-19 and 6-20), which contains at 

most two large displacement amplitude cycles. The data for the two sine wave simulations 

suggests that much, but not all, of the force reduction is recovered between tests. 

Related to these observations, predicting the change in characteristic strength of the lead plug 

over the history of the motion is of interest. Such effects have not been incorporated into our 

numerical simulation models, discussed in Chapters 7-9, but the following is noteworthy. 
114

Kalpakidis and Constantinou (2009a) developed theoretical equations to compute the 

instantaneous strength and evolution of strength in the lead plug over the history of the 

excitation. Kalpakidis et al. (2010) presented a numerical algorithm to incorporate the 

temperature dependence into a bidirectionally coupled bilinear numerical model of the isolator. 

The algorithm to compute the temperature increase in a response simulation is based on the 

dimensions of the lead plug, density and specific heat of lead, thermal diffusity and thermal 

conductivity of steel, and a few other calibrated parameters. To represent the simulation data 

from the series of tests discussed in this report, we have the added challenge of predicting the 

temperature at the start of a simulation, which normally would not be necessary. Incorporating 

heating of the lead plug into numerical simulation models for improved prediction of the data is a 

goal of our ongoing investigation.

Figure 6-19 X-direction force vs. displacement (hysteresis loops) for the 1st and 2nd

simulation of Vogtle 75%
115

Figure 6-20 Y-direction force vs. displacement (hysteresis loops) for the 1st and 2nd

simulation of Vogtle 75%
116

Figure 6-21 X-direction displacement and force histories for the 1st and 2nd simulation of 

Vogtle 75%
117

Figure 6-22 Y-direction displacement and force histories for the 1st and 2nd simulation of 

Vogtle 75%
118

Figure 6-23 X vs. y-direction displacement (displacement trace) for the 1st and 2nd

simulation of Vogtle 75%

-20

-15

-10

-5

0

5

10

15

20

-X Disp (cm)

19.6 21

E

16.3

17.6

S

-20 -10 0 10 20

-20

-15

-10

-5

0

5

10

15

20

18.7

20.2

N

-X Disp (cm)

Y Disp (cm)

-20 -10 0 10 20

16.3

17.6

W

Y Disp (cm)

VOG75-1

VOG75-2
119

Figure 6-24 Y-direction force vs. displacement (hysteresis loops) identified by cycle for 

the 1st and 2nd simulation of Sine 100% (Y)

121

7. MODEL FOR ISOLATION BEARINGS AND BEARING 

CHARACTERIZATION

For further validation, a 3D frame model of the testbed building was developed for numerical 

simulation in the Open System for Earthquake Engineering Simulation v2.2.2 (OpenSees, 

2010). The results of numerical simulation using this model are directly compared to 

experimental data in Chapters 9-10. In this chapter, we describe the modeling and 

characterization of the hybrid LR isolation system. The modeling assumptions for the LR 

bearings and CL bearings are described in Section 7.1 and 7.2, respectively. Section 7.3 

estimates the vertical dynamic properties of the system considering the contribution of the load 

cell assemblies. Section 7.4 details the procedure used by Dynamic Isolation Systems to 

characterize the parameters of the LR bearings based on cyclic testing (Sec. 7.4.1), which is 

extended to the test data (Sec. 7.4.2).

7.1 Lead-Rubber Bearings

The force-deformation relation of the LR bearings was represented by a combination of spring 

elements. The horizontal and vertical behavior of the bearing was assumed to be uncoupled. 

Horizontal-vertical coupling, loss of lateral stiffness, and loss of axial load carrying capacity have 

been observed in elastomeric bearings under the combination of large horizontal displacements 

and axial forces (Buckle and Liu 1994, Buckle et al. 2002, Warn and Whittaker 2006). However, 

the constraint provided by the vertically stiff CL bearings and base diaphragm, documented in 

Chapter 6, prevented axial shortening and loss of lateral load carrying capacity of the isolation 

system, which supports the choice of an uncoupled bearing model. 

7.1.1 Horizontal Direction Modeling Assumptions

The force-displacement relation of the LR bearings in the horizontal direction was idealized as 

bilinear as shown in Figure 7-1. Numerically, this was implemented as a rate-independent 

plasticity model with kinematic hardening; defined by the elastic stiffness (K1), yield force (Fy), 

and the kinematic (KH) hardening modulus. These parameters were calculated from the post 

yield stiffness (Kd) and characteristic strength (Qd) according to: 

1 10 K K = d (7.1)

F Q KD y d dy = + (7.2)

1

1

d

H

d

K K K

K K = − (7.3)

where uy Dy is the yield displacement. The values of Kd and Qd selected for numerical 

simulation are given in Section 7.4. Bidirectional coupling was accounted for through a circular 

yield surface. Physically, the element was implemented as a zero Length Section element in 

OpenSees (2010) incorporating a Bidirectional section to directly represent the forcedeformation relation of the isolator.
122

 

Figure 7-1 Horizontal force-displacement of LR bearing for numerical simulation

7.1.2 Vertical Direction Modeling Assumptions

In the vertical direction, the force-displacement relation of the LR bearings was idealized as 

bilinear elastic with different stiffness in tension and compression as recommended by Dynamic 

Isolation Systems. Physically, the bearings have nominal stiffness in tension but cavitate at 

relatively low values of tensile pressure (Constantinou et al. 2007). This behavior can be 

approximately simulated by assuming a low tensile stiffness. Numerically, an elastic-no tension 

model (Figure 7-2(a)) was combined in parallel with an elastic model (Figure 7-2(b)) to achieve 

the desired behavior shown in Figure 7-2(c). An effective tension stiffness was estimated to be

of 2% of the compression stiffness was assumed. This approach was recommended by 

Dynamic Isolation Systems as an approximate way to capture the elastic-plastic tension 

behavior of the bearing and limited tensile capacity. 

Figure 7-2 Vertical force-deformation of LR bearing for analytical model: (a) elastic-no 

tension, (b) elastic, and (c) combined

The nominal vertical stiffness of each LR bearing, as provided by DIS Dynamic Isolation 

Systems, was Kv = 15,000 kN/cm (8,570 kip/in). The vertical stiffness of the bearings in the 

numerical simulation model was adjusted by trial and error to 10,000 kN/cm (5,710 kip/in) to 

indirectly account for the flexibility of the load cell support assemblies described in Section 4.1. 

This adjusted value was supported by the following approximate calculations. The vertical 

stiffness of a single Type A load cell – present under three of the four bearings – was given as 

KLC = 85,000 kN/cm (48,600 kip/in). The vertical stiffness of the steel plates was assumed to be 

+ =

 (a) (b) (c)
123

dominated by plate bending as the weight carried by the isolator was shifted to different 

locations on the steel plate. Assuming the plate acts as a continuous beam spanning several 

load cells, the plate bending stiffness was computed assuming fixed-fixed boundary conditions 

with a point load (the weight transferred through the isolator) acting midway between the 

supports (load cells):

3

192

plate

EI K

L = (7.4)

where E = 200 GPa (29,000 ksi) is the elastic modulus of steel, L = 75 cm (30 in) is the clear 

length between adjacent load cells, and I = bh3

/12, using b = L/2 and h = 10 cm (4 in) as the 

plate thickness. With these assumptions, the plate bending stiffness was computed to be Kplate = 

28,000 kN/cm (19,000 kip/in), Combining the stiffness of an LR bearing, plate, and load cell in 

series

111 1

K KK K total v plate LC

=+ +

(7.5)

the total vertical stiffness of the bearing load cell assembly was Ktotal = 8,800 kN/cm (5,030 

kip/in), which is close to the assumed value of 10,000 kN/cm (5,710 kip/in). The estimated 

stiffness of the bearing-load cell assembly was estimated similarly for a TP bearing-load cell 

assembly, which was in fact further corroborated by a detailed finite element analysis of the 

assembly (Dao 2012).

7.2 Cross Linear Bearing

Similar to the LR bearings, the force-displacement relation of the CL bearings was represented 

by a combination of spring elements, which were uncoupled in the horizontal and vertical 

directions. This assumption is not strictly accurate since the CL bearing is a friction device, and 

thus the horizontal force is proportional to the instantaneous axial force. However, the friction 

coefficient of the CL bearings was small so that their contribution to the overall base shear was 

negligible. 

7.2.1 Horizontal Direction Modeling Assumptions

The force-displacement relation for the CL bearings was assumed to be elastic-perfectly plastic 

(with a post-yield stiffness of zero) in each horizontal direction as shown in Figure 7-3. The 

sliding rail system acts independently in each horizontal direction; thus a model with a square 

interaction surface was used instead of a bidirectionally coupled model. This assumption only 

affects the first yield mechanism since the model is perfectly plastic. The model initial stiffness 

of the CL bearing (K1s), where yield represents sliding of the bearing, was estimated as:
124

Figure 7-3 Horizontal force-displacement of CL bearing for numerical simulation

1s

y

W K

D

μ= (7.6)

where μ is the coefficient of friction, W is the weight (or static vertical force) on each isolator, 

and Dy is the yield displacement. The assumed friction coefficient for numerical simulation was 

0.0025. Note that this differs from the value listed in Table 3-3, which reflects more recently 

acquired information about the CL bearing. 

Under typical distribution of dead load based on tributary area, the center CL bearing would 

carry more weight than the CL bearings in corner positions. However, in the tested TP isolation 

system, the center bearing was lightly loaded compared to several of the other bearings (Dao, 

2012). The warping at the base of the structure and the shimming procedure used to adjust the 

loads in the LR bearings affected the load distribution. The static vertical force in individual CL 

bearings was not measured during the test program. Since load distribution by tributary area 

was not a reasonable assumption, the total weight carried by the CL bearings (deduced from 

the weight of the structure and the measured weight on the LR bearings) was distributed evenly 

to individual CL bearings according to: 

( )

5

W W total LRBs W − = ∑ (7.7)

Because of the uncertainty in the static axial loads in the CL bearings and the sensitivity of load 

transfer (Section 6.2) to the static equilibrium state, the developed numerical simulation model 

was not expected to accurately track axial forces in either LR or CL bearings. Trial and error 

variation of the vertical stiffness of these devices, which affects the balance of forces and load 

distribution over the isolators, was found to have little consequence to the predicted horizontal 

response of the isolation system.

7.2.2 Vertical Direction

The vertical force-deformation for the CL bearings was modeled using a similar parallel spring 

model as used for the LR bearing. In this case, the compression (Kvc) and tension (Kvt) stiffness 

were independent values provided by the manufacturer as listed in Table 3-3. Thus, the 

composite force-displacement relation was derived from an elastic-no tension spring with 

stiffness Kvc – Kvt (Figure 7.4(a)) and an elastic spring with stiffness Kvt (Figure 7.4(b)) to get the 

combined behavior of Figure 7.4(c). 

Force

Disp

K

1s

Fy = μW
125

Figure 7-4 Vertical force-deformation of CL bearing for analytical model: (a) elastic-no 

tension, (b) elastic, and (c) combined

7.3 Composite Vertical Properties of the Isolation System

As part of our investigation, we considered the possibility that due to the flexibility of the load 

cell assemblies, the vertical stiffness (and fundamental frequency) did not represent a typical 

isolation system in the vertical direction. The following calculations support the conclusion that 

the hybrid LR isolation system was not uncharacteristically flexible in the vertical direction.

A typical “rigid body” vertical frequency of an elastomeric isolation system, computed from

,

v

z rigid

g K

W

ω ⋅ = ∑ (7.8)

ranges from 10-15 Hz (Kasalanati 2012). Recall that the measured weight of the testbed 

building was 5,220 kN (1,174 kip), and the nominal (manufacturer supplied) vertical stiffnesses 

were 15,000 kN/cm (8,565 kip/in) for an LR bearing and 34,700 kN/cm (19,814 kip/in) for a CL 

bearing. The adjusted vertical stiffness of the LR bearing-load cell assembly was assumed to be 

10,000 kN/cm (5,710 kip/in). Ignoring the influence of the CL bearings, suppose the isolation 

system had consisted of 4 LR bearings, which would be typical for the composite weight of the 

system, and not supported on load cells - thus representative of the expected field conditions for 

these LR bearings. The vertical frequency of the isolation system, computed from Equation 

(7.8), would be about 17 Hz (0.06 sec), which is on the stiff side of typical. Now, suppose the 

isolation system consisted of 4 LR bearings supported on load cells with the modified stiffness 

of 10,000 kN/cm (5,710 kip/in). In this case, the frequency would be reduced slightly to 14 Hz 

(0.07 sec), which is also stiff. However, the actual hybrid system tested in this experimental 

program, with 4 LR bearings on the load cell assemblies at 10,000 kN/mm (5,710 kip/in) and 5 

CL bearings at 34,700 kN/cm (19,814 kip/in), had a vertical frequency of 31 Hz (0.032 sec). 

Thus, as asserted above, the fundamental vertical frequency of the tested hybrid isolation

system was actually quite large, such that the system can be considered sufficiently stiff.

7.4 Characterization of Lead-Rubber Bearings

To predict the response of the isolation system for design of a nuclear power plant, one would 

ideally develop the modeling or numerical simulation parameters based on physical properties 

of the individual isolation devices. As discussed in Chapter 1, Kd is physically related to the 

stiffness of rubber and Qd is physically related to the strength of the lead plug. For design, a 

single set of bilinear parameters (Kd, Qd) is generally used to represent the bearing hysteresis 

loop over a wide range of displacement, supplemented by bounding analysis. 

Kvc

Kvt Kvt

Kvc - Kvt

+ =

 (a) (b) (c) 
126

For the LR bearings tested as part of this experimental program, a single set of simulation 

parameters did not lead to sufficient accuracy in the model over the range of displacement 

amplitudes observed in the test program. Factors that contributed to the disparity in bearing 

hysteresis loop and best fit model parameters included the following. First, the test program 

included a few small amplitude simulations, such as Westmorland 80% that did not drive the 

isolators sufficiently into the nonlinear range to develop the full characteristic strength of the 

lead plug. Smaller amplitude motions would not normally be considered in design. Second, the 

pinching behavior induced by the smaller size lead plug, not seen in typical full scale LR 

bearings, meant that a bilinear model could not be fit closely to the observed hysteresis loop,

which lacked a consistent backbone curve. Thus, more significant parameter variations induced 

by amplitude changes were observed throughout the test program. Finally, several a couple

sources of characteristic strength variation were not incorporated into the bilinear model. For 

one, lack of a full recovery time between simulations meant that the response of the bearings 

was affected by temperature variation (as documented in Section 6.3), which could not be 

incorporated into the model easily due to the absence of measured temperature data. 

Furthermore, the response of the bearings was not representative of the real (unheated) state 

that would be experienced for a real earthquake when the isolation system is not excited in 

regular intervals. Second, the effective yield stress of the lead plug is rate dependent, wherein 

rate variation is a function of amplitude variation. Third, axial force variation affected the 

confinement of the lead plug, a dynamic effect that also was not accounted for in the bilinear 

model.

To obtain a consistently accurate prediction of isolator displacements and forces across the set 

of trials, we elected to characterize the bearing parameters independently for each simulation in 

the test program. As described above, this procedure was necessary only due to the special 

circumstances of the test program, and would not nor could not be attempted for design. 

Characterized bearing properties were determined both for pseudo-static cyclic tests conducted 

by DIS and for each simulation during the test program at E-Defense.

7.4.1 Characterization by Dynamic Isolation Systems (DIS) 

The LR bearings were characterized by Dynamic Isolation Systems in their manufacturing 

facility prior to shipment to Japan. Bearings were tested in the machine two at a time; each pair 

of bearings was subjected to cyclic shear tests under constant compressive load and the 

measured horizontal force represents the sum of the horizontal forces in the two bearings. The 

tests were displacement controlled, such that the bearings were cycled back and forth to the 

target maximum displacement in each direction for the desired number of cycles. A series of 

four tests were conducted at different axial loads and displacements, as summarized in Table 7-

1. Test C was added to accommodate E-Defense’s safety protocol, given that a displacement 

demand of 55 cm (22 in) was to be targeted during the tests. Test D was a repeat of Test A and 

was intended to document any change in hysteresis loops as a result of repeated loading. A 

minimum fifteen minute interval was inserted after every test. As shown by the rate parameter in 

Table 7-1, these characterization tests were essentially static and thus do not include any rate 

effects on the bearing response. In particular, the lead plug heating effects would be smaller in a 

static simulation that in a high speed cyclic simulation.
127

Table 7-1 Compression Shear Test Schedule. 

Test ID Number of 

Cycles

Axial Load

(kN)

Dmax

(cm)

Shear Strain

 % 

Rate 

(cycles/min)

A 3 600 30 125 1.15

B 3 1000 50 208 0.71

C 0.5 100 65 271 0.54

D 3 600 30 125 1.15

In the test report provided by Dynamic Isolation Systems (Appendix E), the isolator properties 

were determined by fitting a bilinear loop to the recorded hysteretic loop such that the energy 

dissipated and the effective stiffness of the two loops were equal. The fitted second-slope postyield stiffness (Kd) and characteristic strength (Qd) were determined directly from the fitted loop. 

The fitting procedure is described conceptually by Figure 7-5. The effective stiffness of the 

isolator (Keff) is equal to Fmax/Dmax, where Dmax is the maximum isolator displacement and Fmax is 

the maximum force measured in the isolator. If the cycle is unsymmetric, the peak-to-peak 

stiffness is used rather than the half cycle effective stiffness. The energy dissipated per cycle 

(EDC = area of the loop) was determined by numerically integrating the force-displacement 

data. Fixing the corner points (Fmax, Dmax and Fmin, Dmin, which are the minimum isolator force 

and displacement, respectively) of the analytical model to match the test data, Qd and Kd are 

adjusted until the energy dissipated in a cycle of the theoretical bilinear loop with initial stiffness 

K1 = 10 Kd matches the numerically integrated energy dissipated from the recorded data. 

Figure 7-5 Comparison of fitted hysteresis loop and parameters to test data

The force-displacement relationship for one of the isolator pairs recorded during the Test A 

loading protocol is shown in Figure 7-6. Pinching of the hysteresis loop is observed around zero 

displacement, which is expected when the lead plug is small relative to the diameter of the 

bearing or simply small on an absolute scale. Pinching may be observed in full size or prototype 

LR bearings manufactured for real world projects, but far less pronounced than that observed 

here (Kasalanati, 2012). 

The recorded and fitted parameters for the test data (Figure 7-6) are listed in Table 7-2 for each 

cycle as well as the average over all 3 cycles. The energy dissipation per cycle EDC and thus 

Displacement

Kd

Force

Fmax

Dmax

Qd
128

fitted Kd and Qd decrease after each cycle, with a large drop noted after the first cycle. As a 

result of the pinching, the fitted Qd is significantly larger (by up to 70%) than the y-axis force 

intercept of the test data. The characterization procedure was carried out for each cycle of all 

four tests. The characterized values for each test (averaged over all cycles and over the two 

pairs of bearings) are reported in Table 7-3.

 

Figure 7-6 Hysteresis loop of Test A and D for one of the isolation pairs

Table 7-2 Recorded and fitted parameters for Test A. 

Cycle Dmax

(cm)

Fmax

(kN)

Keff

(kN/cm)

EDC

(kN.cm)

Kd

(kN/cm)

Qd

(kN)

1 30.1 266 8.8 8455 6.5 71.2

2 30.2 249 8.2 7486 6.2 62.9

3 30.2 245 8.1 7206 6.1 60.4

AVERAGE 30.2 253 8.4 7716 6.3 64.8

Table 7-3 Characterized isolator parameters for all tests in the sequence. 

Test Axial Load

(kN)

Dmax

(cm)

Kd

(kN/cm)

Qd

(kN)

Keff

(kN/cm)

A 600 30 6.3 64.7 8.4

B 1000 50 5.1 75.6 6.6

C 100 65 5.5 85.6 6.8

D 600 30 5.8 63.1 7.9

-30 -15 0 15 30

-300

-200

-100

0

100

200

300

Test A

Displacement (cm)

Force (kN)

-30 -15 0 15 30

Test D

Displacement (cm)
129

The isolator parameters given in the design specifications (Table 3-3) were stiffness Kd = 6.5 

kN/cm (3.7 kip/in), characteristic strength Qd = 65.7 kN (14.8 kips), and effective stiffness Keff = 

8.7 kN/cm (5 kip/in). The fitted parameters are within 4% of the design specifications at a 

displacement of 30 cm (12 in). Note that the design specifications are just target values set by 

Dynamic Isolation Systems prior to their manufacture.

7.4.2 Characterization Based on Experimental Data

The characterization of the bearings for the earthquake simulations was complicated by the fact 

that the experimental data was bidirectional, and the random earthquake excitation did not 

produce the smooth controlled hysteresis loops of cyclic data. We searched for a procedure to 

characterize the bearing parameters directly based on bidirectional data, but without success. 

Thus, the following alternative procedure was implemented to characterize the bearings for 

each test.

Step1: To obtain the best results for a unidirectional characterization procedure, the test data in 

both directions was rotated to an alternate coordinate system with a main axis that contained 

the largest displacement cycle of the record. The main axis was identified from the largest single 

cycle peak-to-peak excursion on the displacement trace. For instance, in Vogtle 100%, the main 

axis was identified at an approximate 45 degree rotation (Figure 7-7), consistent with the 

rotation of the input motion to generate the peak displacement demand along the diagonal 

(Section 3.3). Identification of the main axis was subjective if the displacement trace contained a 

large circular cycle of motion. 

Figure 7-7 Projection of displacement trace to main axis for Vogtle 100%

Step 2: The isolator displacement history was projected to the main axis direction of the rotated 

data, and the cycle containing the largest peak-to-peak displacement was selected for 

characterization. Figure 7-8 shows the selected cycle for LRB-E for Vogtle 100%. 

Main 

axis
130

Figure 7-8 Selection of cycle for characterization for Vogtle 100%

Step 3: The energy dissipated (EDCtest) for the selected cycle was determined by numerically 

integrating the shear force versus lateral displacement using a cumulative trapezoidal algorithm.

Step 4: The theoretical characteristic strength Qd, post-yield stiffness Kd and yield displacement 

uy were fitted to the projected test data using essentially the same algorithm employed by 

Dynamic Isolation Systems, and summarized as follows. The energy dissipated in a bilinear 

force-displacement loop EDCbilin is determined by:

EEEEEEbbbbbbbbbb = 4QQdd(uu pppp pp − yy) (7.9)

where Qd is the characteristic strength, upeak Dpeak is absolute peak displacement for the selected 

cycle, and Duy is the yield displacement. The energy dissipated during the experiment (EDCtest) 

is equated to the theoretical energy dissipated (EDCbilin), and Equation 7.9 is rearranged to 

solve for Qd. The yield displacement uyDy, which is unknown, is dropped from the equation and 

replaced with a calibration factor γβ initialized to 1:

QQdd = γγγγ∗ tttttttt

4∗DDDDpp pp

 (7.10)

From the estimate of Qd, the theoretical second-slope post-yield stiffness Kd and yield 

displacement uy Dy are computed.

KKdd = FFpp pp−QQdd

pp pp

 (7.11)

yy = QQdd

KK1−KKdd

 (7.12)

5 10 15 20 25

-25

-20

-15

-10

-5

0

5

10

15

20

25

Time (sec)

Displacement (cm)

 

X

Y

Proj.

Selected Cycle
131

where, Fpeak is the absolute peak displacement force for the selected cycle, and K1 = 10 Kd. 

Equations (7.9) – (7.12) are computed iteratively until convergence is obtained (EDCbilin ≈ 

EDCtest). If EDCtest < EDCbilin, then γβ is decreased by small increments, while if EDCtest > 

EDCbilin, then γβ is decreased increased by small increments. 

Once the characterized values were computed by the above procedure, the model parameters 

(initial stiffness, kinematic hardening modulus and yield force) were computed from Equations 

7.1 to 7.3. 

The hysteresis loop for the projected experimental data and the fitted loop based on 

characterized parameters are compared in Figure 7-9 for Vogtle 100%. In Figure 7-9, no 

obvious bidirectional interaction is observed in the experimental data, which supports the idea 

that projecting the force-displacement data to a main axis improves the characterization 

compared to experimental data that has not been projected. 

Figure 7-9 Hysteresis loop of peak cycle for the projected-direction for East bearing for 

Vogtle 100%

The characterized parameters Qd, Kd, uy Dy for each bearing and the average among all 

bearings for each simulation are summarized in Table 7-4. The numerical simulation model in 

OpenSees used the average values listed on the last column for each simulation, with some 

minor adjustment, which is described momentarily. The range of the parameters varied as 

follows. The average of Qd ranged from 33.4 – 89.4 kN (7 - 20 kip), and the average Kd ranged 

from 5.7 - 11 kN/cm (3.2 - 6.3 kip/in). However, omitting WSM80 and ELC130, which produced 

only about 9 cm (4 in) and 21 cm (8 in) displacement, respectively, Qd ranged from 62.0 - 89.4 

kN (14.0 – 20.1 kip), and Kd ranged from 5.7 - 8.2 kN/cm (3.2 – 4.7 kip/in), where the 

displacement varied from 20 cm (8 in) to 55 cm (22 in). As a comparison, the target 

specifications at a displacement of 30 cm (12 in) were Qd = 65.7 kN (14.8 kip) and Kd = 6.5 

kN/cm (3.7 kip/in).

-30 -20 -10 0 10 20 30

-300

-200

-100

0

100

200

300

Displacement (cm)

Force (kN)

Experimental

Fitted
132

The numerical and experimental simulation hysteresis loops are compared using the design 

parameters (Figure 7-10(a),(c)) and the parameters determined by characterization (Figure 7-

10(b),(d)) for the Westmorland 80% and Diablo 95% (XY) motion. The numerical simulation 

loops were determined by response history analysis of the complete specimen model 

(described in Chapter 8) subjected to the recorded table motion. The characterized parameters 

led to a clear improvement in prediction of the peak displacement for Westmorland 80%. On the 

other hand, since the characterized and design values of Qd and Kd for Diablo 95% (XY) are 

about the same, the hysteresis loops and peak displacements generated by the two approaches 

were similar, as shown in Figure 7-10(c) and (d). Overall, Figure 7-10 confirms that the bilinear 

model is a reasonable assumption for the behavior of the bearing if the parameters are 

calibrated for each simulation. 

Table 7-4 Characterized bearing parameters for each earthquake simulation. 

Trial # GM Isolator Qd

(kN)

Kd

(kN/cm)

K1

(kN/cm)

uy Dy

(cm) Average

1 WSM80

E 37.1 10.7 106.6 0.4 Fy = 37.1 kN

S 31.5 10.7 107.0 0.3 Qd = 33.4 kN

N 33.6 11.0 110.3 0.3 Kd = 11.0 kN/cm

W 31.5 11.5 115.1 0.3

2 SIN100(Y)-1 

E 78.2 8.4 83.8 1.0 Fy = 86.0 kN

S 76.3 8.1 80.8 1.0 Qd = 77.4 kN

N 77.8 8.2 81.6 1.1 Kd = 8.2 kN/cm

W 77.2 8.0 80.4 1.1

3 VOG75-1 

E 77.4 8.2 81.9 1.0 Fy = 75.5 kN

S 61.6 8.7 87.2 0.8 Qd = 67.9 kN

N 70.8 8.2 81.8 1.0 Kd = 8.2 kN/cm

W 61.8 7.6 75.9 0.9

4 VOG100

E 86.3 7.4 73.5 1.3 Fy = 88.6 kN

S 77.6 7.7 76.7 1.1 Qd = 79.7 kN

N 82.0 7.2 71.5 1.3 Kd = 7.4 kN/cm

W 72.9 7.3 73.5 1.1

5 VOG125

E 91.9 6.4 64.2 1.6 Fy = 95.3 kN

S 83.0 6.4 63.8 1.4 Qd = 85.8 kN

N 86.9 6.4 64.3 1.5 Kd = 6.5 kN/cm

W 81.3 6.7 67.4 1.3

6 VOG150

E 92.4 6.0 59.8 1.7 Fy = 98.5 kN

S 87.4 5.6 55.6 1.7 Qd = 88.7 kN

N 87.5 5.9 59.1 1.6 Kd = 5.9 kN/cm

W 87.3 6.0 59.9 1.6

7 VOG175

E 92.3 6.0 60.0 1.7 Fy = 99.4 kN

S 90.5 5.4 53.8 1.9 Qd = 89.4 kN

N 87.8 5.7 56.7 1.7 Kd = 5.7 kN/cm

W 87.1 5.6 56.1 1.7

8 DIA80

E 79.2 6.6 65.7 1.3 Fy = 74.8 kN

S 64.0 6.3 63.2 1.1 Qd = 67.3 kN

N 68.4 6.9 68.5 1.1 Kd = 6.6 kN/cm

W 57.6 6.7 67.0 1.0
133

Table 7-4 (Cont.) Characterized bearing parameters for each earthquake simulation. 

Trial # GM Isolator Qd 

(kN)

Kd 

(kN/cm)

K1 

(kN/cm)

uy Dy 

(cm) Average

9 DIA95(XY)

E 79.9 6.3 62.8 1.4 Fy = 78.1 kN

S 65.7 6.0 59.9 1.2 Qd = 70.3 kN

N 74.4 6.2 62.4 1.3 Kd = 6.2 kN/cm

W 60.9 6.3 63.0 1.1

10 ELC130

E 61.0 7.7 77.2 0.9 Fy = 59.5 kN

S 56.6 8.1 80.9 0.8 Qd = 53.5 kN

N 48.5 8.8 88.3 0.6 Kd = 8.4 kN/cm

W 47.8 9.0 89.6 0.6

11 IWA(XY)

E 81.3 6.7 66.7 1.4 Fy = 87.6 kN

S 80.7 5.9 58.5 1.5 Qd = 78.8 kN

N 75.1 6.5 65.2 1.3 Kd = 6.3 kN/cm

W 78.0 6.3 62.7 1.4

12 RRS88(XY)

E 75.5 6.6 66.2 1.3 Fy = 75.8 kN

S 65.1 6.4 63.6 1.1 Qd = 68.2 kN

N 68.8 6.7 66.9 1.1 Kd = 6.6 kN/cm

W 63.3 6.6 65.9 1.1

13 RRS88

E 74.3 6.5 65.3 1.3 Fy = 74.7 kN

S 64.5 6.2 62.1 1.2 Qd = 67.2 kN

N 67.8 6.5 65.0 1.2 Kd = 6.4 kN/cm

W 62.2 6.5 64.7 1.1

14 VOG75-2 

E 69.8 7.6 75.9 1.0 Fy = 69.5 kN

S 61.3 7.6 75.7 0.9 Qd = 62.5 kN

N 64.3 7.2 72.5 1.0 Kd = 7.4 kN/cm

W 54.8 7.2 72.4 0.8

15 SIN100(Y)-2 

E 75.5 7.7 76.5 1.1 Fy = 79.7 kN

S 74.6 7.3 72.8 1.1 Qd = 71.8 kN

N 72.3 7.3 73.1 1.1 Kd = 7.4 kN/cm

W 64.6 7.2 71.9 1.0
134

Figure 7-10 Analytical and experimental hysteresis loop using (a),(c) isolator design 

parameters and (b),(d) characterized parameters

The characterized model was calibrated for a cycle close to the peak displacement, but 

consequently did not improve the response prediction for cycles at smaller displacement. For 

example, the y-direction hysteresis loop and force history (test versus analysis) are shown in 

Figure 7-11 for the first sine wave trial. At small amplitudes (after 22 sec in Fig 7-11(b) and (d)), 

the actual force in the LR bearing was quite small, indicating the lead plug was not fully 

engaged. Thus, both numerical models significantly overestimated the force in this range. The 

characterized model gave a much better prediction of peak force than the design model at the 

expense of higher force error in the small amplitude part of the record. As another example, the 

x-direction mean square error in the design model was lower than the characterized model for 

Vogtle 175%. Based on the displacement history (Figure 7-12), the characterized model better 

-100

0

100 E

(a) WSM80% - Design

S E

(b) WSM80% - Characterized

S

-10 0 10

-100

0

100 N

Force X (kN)

-10 0 10

W

Displacement (cm)

-10 0 10

N

-10 0 10

W

Numerical Experimental

-400

-200

0

200

400 E

(c) DIA95%(XY) - Design

S E

(d) DIA95%(XY) - Characterized

S

-50 0 50

-400

-200

0

200

400 N

Force X (kN)

-50 0 50

W

Displacement (cm)

-50 0 50

N

-50 0 50

W
135

predicted the displacement for the largest cycle (around 10 sec), but the design model better 

predicted the displacement in subsequent cycles (compare Figures 7-12(a) and (b)). 

 Figure 7-11 Y-direction hysteresis and force history of test data of the East bearing 

compared to (a)-(b) characterization model, and (c)-(d) design model for 

Sine 100%(Y)-1 

 

The comparisons illustrate the challenge of predicting the bearing response using a single 

bilinear model that is insensitive to the characteristic strength variation due to axial force 

variation amplitude dependence and temperature change in the lead plug. The challenge was 

amplified by the pinching in the lead plug, which is not typical of full scale isolation bearings. 

When the bearing model was calibrated to the largest displacement cycle, it tended to 

overestimate the force and underestimate the displacement at smaller amplitudes. Development 

or use of existing models that account for the various effects such as pinching and, thermal, and 

axial force may significantly improve the fitted response that could be obtained from a model 

with a single set of parameters. One potential improvement was attempted, which was to model 

the bearings with trilinear force-deformation that may be able to represent a smoother transition 

to the fully-yielded state, and thus reduce the energy dissipated in the hysteresis loop for small 

cycles. This potential improvement was attempted, and is described in the next section. 

-50 0 50

-200

0

200

Force Y (kN)

SIN100%(Y)-1 East Bearing

 Characterized 

0 5 10 15 20 25 30

-200

0

200

-40 -20 0 20 40

-200

0

200

 Design 

Force Y (kN)

Disp. Y (cm)

0 5 10 15 20 25 30

-200

0

200

Time (sec)

Numerical Experimental

(a) (b)

(c) (d)
136

Figure 7-12 X and y-direction displacement history of experimental data compared to 

(a) characterized model and (b) design model for Vogtle 175%, East bearing

7.4.3 Trilinear Characterization

The elastic stiffness in a bilinear hysteretic model is determined by the stiffness of the lead plug, 

and the model assumes that the lead plug stiffness is linear. However, the experimental data 

exhibited a smooth transition from the linear to the post-yield state (e.g. Sine 100%(Y)-1, Figure

7-13), such that a numerical model defined by a sharp transition from the linear to post-yield 

state was overly stiff for large displacements. The gradual transition to the yielded state is a 

reflection of the true behavior of the lead plug (Figure 7-14). Thus, a trilinear model was 

attempted to improve the small displacement prediction without altering the large displacement 

response of the model. (A Bouc-Wen model (Park et al. 1986) is also known for smoothing the 

transition from the linear to the post-yield state, and is another approach that could have been 

attempted.)

-50

0

50

Disp. X (cm)

VOG175 - East Bearing 

(a) Characterized

5 10 15 20 25

-50

0

50

Disp. Y (cm)

Time (sec)

-50

0

50

(b) Design

Disp. X (cm)

5 10 15 20 25

-50

0

50

Disp. Y (cm)

Time (sec)

Numerical

Experimental
137

Figure 7-13 Force vs. displacement loop for the East LRB in Sine 100%(Y)-1 

Figure 7-14 Conceptual force vs. displacement of lead plug for monotonic loading 

(courtesy of Dynamic Isolation Systems)

-30 -20 -10 0 10 20 30

-300

-200

-100

0

100

200

300

Disp. Y (cm)

Force Y (kN)

SIN100%(Y)-1 East Bearing

Numerical Experimental

0 1 2 3 4 5 6 7

0

10

20

30

40

50

60

70

Displacement (cm)

Lead Force (kN)

Experimental

Fitted

KL1 = 4

KL2=0.86

KL3=0.1
138

In the trilinear model, the stiffness of the lead-plug was modeled as piecewise linear with 

stiffness KL1, KL2 and KL3. The values of KL1, KL2 and KL3 from Figure 7-14 were selected such 

that the tangent stiffness for the displacement less than 1 cm and greater than 6 cm matched 

the experimental data. Then, KL2 was selected such that the same area under the actual curve 

and the theoretical curve were equal area under the actual curve and the theoretical curve were 

equal. 

The trilinear model was implemented in OpenSees as three bilinear springs in parallel, where 

each spring was assigned different properties and represented bidirectionally coupled behavior. 

Springs 1 and 2 represented the stiffness and energy dissipation in the lead plug while Spring 3 

represented the stiffness of rubber. Springs 1 and 2 were elastic-perfectly plastic with initial 

stiffness and yield force as determined by Figure 7-14. The properties of the third spring 

representing the rubber were calculated as follows. The elastic stiffness of rubber (K1,rub) was 

computed as

KK1,rr = KK1,cchaa − KK1, (7.13)

where K1,char is the characterized elastic stiffness described in Section 7.4 and K1,lead is the 

elastic stiffness from Figure 7-15. The characteristic strength of rubber Qd,rub was defined as:

QQdd,rr = QQdd,cchaa − QQdd, (7.14)

where Qd,char is the characterized characteristic strength described in Section 7.4 and Qd,lead is 

the characteristic strength of the lead from Figure 7-15. The yield displacement was defined as:

yy,rr = QQdd,rrrrrr

KK1,rrrrrr−KKdd,cch

 (7.15)

where Kd,char is the characterized post yield stiffness described in Section 7.4.2. These 

properties were sufficient to determine the modeling parameters of the spring. 

Results for numerical and experimental simulation with both the bilinear and trilinear bearing 

models are compared in Figure 7-15 for SIN 100(Y)-1 (Figure 7-15(a)-(b)) and Vogtle 100% 

(Figure 7-15(c)-(d)), where the influence of the multi-linear pivoting can be seen especially in the 

center of the loop and the large displacement transitions for Sine 100%. The trilinear model led 

to some improvement in the small displacement hysteresis for SIN100(Y)-1, but the 

improvement was negligible for Vogtle 100%. The same was true for the other earthquake 

simulations. In particular, the trilinear model did not appear to substantially change the 

prediction of the displacement over the majority of the record, and did not address the primary 

inconsistency of the model compared to the experimental data, which was the pinching of the 

lead plug through the center of the loop that extended into the post-yield behavior. Therefore, 

the trilinear model was not adopted for final numerical simulation to validate the experimental 

data.
139

Figure 7-15 Comparison of numerical and experimental simulation force versus 

displacement of the East bearing for SIN100(Y)-1 and VOG100 using (a), (c) a 

trilinear hysteretic model and (b), (d) a bilinear hysteretic model

7.4.4 Final Bearing Parameter Selection for Numerical Simulation

One slight modification was made to the characterized bearing parameters used in the final 

numerical simulation models, which are compared to the experimental results in Chapter 9. The 

peak displacements using the best fit parameters of Table 7-4 were often below the peak 

experimentally observed displacements observed. The problem seemed to result from the fact 

that the fitted Qd (y-intercept) was larger than the actual Qd of the LR bearings, thus increasing 

the energy dissipation at lower amplitudes and suppressing the higher displacement amplitudes 

from ever being reached. By trial and error, we observed that decreasing Qd by a nominal 

amount relative to the best fit value improved the displacement prediction. As an example, 

Figure 7-16 compares simulation results using the best fit characterized model parameters and 

the model with Qd reduced to 85% of the characterized value to the experimental data for Vogtle 

100%. The model with reduced Qd better predicted the peak displacement amplitude and 

followed the bearing hysteresis better than the best fit characterized model. Thus, the reduced 

value of Qd was used for all simulation results presented in Chapter 9.

-300

-200

-100

0

100

200

300

Force Y (kN)

(a) SIN100(Y)-1 - Trilinear model (b) SIN100(Y)-1 - Bilinear model

-20 -10 0 10 20

-300

-200

-100

0

100

200

300

(c) VOG100 - Trilinear model

Force X (kN)

Displacement Y (cm)

-20 -10 0 10 20

(d) VOG100 - Bilinear model

Displacement Y (cm)

Numerical

Experimental
140

Figure 7-16 Comparison of numerical simulation with 100% and 85% characterized 

values of Qd to the experimental data for Vogtle 100%; displacement history 

and hysteresis loop in the y-direction for the East bearing

8 10 12 14 16 18 20

-20

-10

0

10

20

East Bearing Displacement History - Y

Displacement (cm)

Time (sec)

-20 -10 0 10 20

-300

-200

-100

0

100

200

East Bearing Force vs Disp. - Y

Force (kN)

Displacement (cm)

100% Qd

85% Qd

Test
141

8. NUMERICAL SIMULATION MODEL OF TESTBED STRUCTURE

As introduced in Chapter 7, a model for numerical simulation of the testbed building with the 

isolation system was developed in OpenSees. The assumptions used in developing the model 

of the building frame with floor slabs, shown in Figure 8-1(b), are described in this chapter. As 

an overview, the contribution of floor slabs to the bending stiffness was included in composite 

beam sections and their in-plane stiffness was accounted for through application of diaphragm 

constraints. The beam-column connection behavior was represented by a panel zone model. 

Material nonlinearity was considered through nonlinear material models, and some geometric 

nonlinearities were included through a P-Delta transformation. Mass and gravity loads were 

lumped to nodes. Rayleigh damping was applied to the superstructure with additional inter-story 

dampers to represent energy dissipation. The bearings were represented using the modeling 

assumptions and characterized parameters of LR bearings and CL bearings described in 

Sections 7.1 and 7.2, with calibrated bearing parameters of Table 7-4 Chapter 7. 

Figure 8-1 (a) Photo and (b) OpenSees model of the testbed building without isolators

(a) (b)
142

Figure 8-2 Primary beam, beam-to-column connection, and slab

8.1 Modeling Beams and Columns

The beams of the testbed building were composed of either rolled or built-up I-sections. Primary 

beams - supported by columns - consisted of a small section segment in the middle bolted to 

large section segments at the ends, all 40 cm (16 in) deep (Figure 8-2). The beam-column 

connections were fully restrained moment connections with beam flanges and webs welded to 

the column face. Generally, the primary beams were haunched at the ends for improved 

strength, and continuity plates protected the panel zones. Secondary beams - supported by 

primary beams - were connected to the primary beams through shear tabs. The columns were 

made of 35 cm (14 in) HSS sections with thickness varying from story to story (Appendix A). 

Primary beams were modeled by displacement-based nonlinear frame elements with distributed 

plasticity. The displacement formulation was selected over advocated force formulations 

(Neuenhofer and Filippou 1997) to improve the convergence of the analytical model. To 

optimize the performance of displacement-based elements, each beam member was divided 

into at least 8 elements. Since mass was lumped at the nodes, the discretization also helped to 

distribute mass over the structure more realistically.

The nonlinear frame elements were accompanied by composite section models to account for 

the contribution of floor slabs, which were connected to the primary beams through shear studs 

(Figure 8-2). The effective slab width for each side of the composite section was the minimum of 

(1) one-eighth of the beam span, (2) one-half the distance between the beams and (3) the 

distance to the edge of the slab (AISC 2005). Longitudinal slab reinforcement was included in 

the section model. The Giuffre-Menegotto-Pinto (CEB, 1996) and Kent-Park concrete materials 

(Kent and Park 1971, Scott et al. 1982) are used to model the stress-strain relations of steel and 

unconfined concrete (applied to floor slabs), respectively. The tensile resistance of concrete was 

neglected in the concrete material model.

Corrugated deck slab

Concrete protecting 

shear studs

Connection between 

segments

Large section 

segment

Stress 

haunch

Continuity 

plate

Small section 

segment
143

For these non-symmetric composite sections, when the material behavior becomes nonlinear, 

the neutral plane of the section moves and the geometric centerline deforms axially under pure 

bending loads. However, the rigid diaphragm constraint prevents the axial deformation of the 

centerline, thus introducing an axial force to the bent beam. The axial force changes the 

behavior of beams significantly, as demonstrated in Figure 8-3(b), where the bending behavior 

with and without axial deformation restraint are compared for a simply supported 5 m (16 ft) 

beam element with a composite section driven cyclically at the midpoint. To avoid the 

unintended effect of axial force on bending of the composite beam sections, the axial and 

bending behaviors were decoupled through the use of resultant section models for momentcurvature and axial force-strain. The resultant beam section behavior was determined from 

analysis of the composite fiber sections (e.g. Figure 8-3(a)). As an example, the pure bending 

cyclic behavior of the composite fiber section determined by section analysis (solid line in Figure 

8-4(a)) was approximately represented by combining the steel material model (Figure 8-4(b)) 

with a hysteresis model (Figure 8-4(c)) in parallel.

Figure 8-3 Behavior of a representative composite fiber section beam with and without 

axial restraint: (a) fiber section geometry and (b) force-displacement 

relationship

Secondary beams of the testbed were modeled as elastic beam elements with elastic composite 

sections. The secondary beams in the model were also divided into 8 elements to distribute 

mass realistically.

The testbed columns were also modeled by displacement-based nonlinear frame elements, but 

fiber sections were used to account for axial force-bending interaction in the columns. Each 

column member was modeled with 3 elements to improve the performance of the displacementbased elements. Because plasticity is mainly concentrated at the two ends of a column 

member, the length of the end elements was set equal to the section height (35 cm or 14 in). 

Three displacement-based elements were shown to give similar moment rotation behavior to 

the force-based element with 7 integration points (Dao 2012). 

During the experimental simulation, the response of the testbed frame was essentially elastic. 

However, the floor system exhibited some nonlinearity due to the cyclic response of concrete 

between tension and compression. Dao et al. (2012b) demonstrated that the numerical 

simulation results more closely matched the experimental data when nonlinear modeling 

assumptions were considered, compared to an elastic frame model. Thus, we believe that the 

attention to detail and consideration of nonlinear effects in the model was justified. 

-4 -2 0 2 4

-1

-0.5

0

0.5

1

Displacement, ∆ (cm)

Force, P (MN)

Axially restrained

Axially unrestrained

Reinforcement (b)

Concrete slab

1250 × 116 mmmm

Steel I-section

400 × 200 × 9 × 19 mmmm

(a)
144

Figure 8-4 Composite section behavior: (a) moment-curvature relationship of the 

section, (b) component 1 of resultant section modeled by steel material 

model, and (c) component 2 of resultant section modeled by hysteresis 

model

8.2 Modeling Panel Zones

The Krawinkler panel zone model (Krawinkler 1978, Charney and Downs 2004) was used to 

model the connection between beams and columns. According to this model, each panel zone 

(Figure 8-5(a)) was modeled by 8 rigid elements and 2 elastic-perfectly plastic rotational 

springs, one representing the shear behavior of the panel zone (or the web, lying in the working 

plane) and one representing the bending behavior of the flanges (perpendicular to the working 

plane) (Figure 8-5(b)). Since the columns were fully welded to primary beams in both 

directions, the panel zones in two directions were independently modeled by two Krawinkler 

panel zone models.

(c)

(a)

(b)
145

Figure 8-5 Panel zone model for beam to column connection. (a) beam to column 

connection, (b) numerical model of panel zone

The initial stiffness SSPP and yield strength MMyy of the spring representing the shear behavior of 

the panel zone web were computed as:

SSPP = GGVVPP (8.1)

MMYY = 0.58FFYYVVPP (8.2)

where: GG = shear modulus of steel, VVPP = volume of the panel zone web, and FFFFYY = yield 

strength of steel material. The initial stiffness SSFF and yield strength MMyy of the spring 

representing bending of the flanges are:

SSFF = 0.75GGbb tt2 (8.3)

MMYY = 1.80FFYYbb tt2 (8.4)

where: bb = flange width of column and tt = flange thickness of column.

Elastic axial elements equivalent to the Krawinkler model were also used to model gusset plates 

(Figure 8-6(a)), which were an integral part of the specimen for attaching dampers in the March 

2009 test (Kasai et al., 2010). The dampers were not present during the experimental simulation 

described in this report. Finite element analysis of a connection with gusset plate subjected to 

gravity load suggested that the gusset resistance is in the diagonal direction (Figure 8-6(b)), and 

can be modeled as a diagonal strut. The equivalent elastic stiffness kktt tttt of the axial element 

was:

 kktt tttt = SSPP

�xx2+yy2�

xx2yy2 (8.5)

where xx × yy = size of the gusset plate and SSPP = initial shear stiffness of the gusset plate 

(Equation (8.1)).

Panel web

Beam Column

Beam

Column

Rigid element

Hinge

Spring 

representing 

column flanges

Spring 

representing 

panel web

(a) (b)
146

Figure 8-6 Gusset plate and its finite element model: (a) gusset plate, (b) Von-Mises 

stress due to gravity load

8.3 Modeling Gravity Load and Mass

As mentioned earlier, the testbed building was modeled as a bare frame without slabs so that 

gravity loads and mass were applied directly to beams and columns. Static analysis of a 

SAP2000 model subjected to gravity loads with slabs represented by general shell elements 

was used to compute beam internal forces and distributed loads. From the shear forces VV and 

VVjj at the two ends of a beam element, the equivalent uniform load vv on the beam element was 

computed according to:

 vv = VVii−VVjj

LL (8.6)

where VV and VVjj are shear forces at the two ends of the element, and LL is the length of the 

element. The mass of the OpenSees model was directly derived from the computed distributed 

loads and lumped to every node of the model. Because of the bending stiffness of the slab, 

some of the slab gravity load transferred directly to the corner slab nodes. In the OpenSees 

model, these loads were applied as concentrated loads to the corresponding corners. 

Table 8-1 summarizes the weight and the eccentricity of gravity center from the geometric 

center of the SAP and OpenSees models. The weight of all floors in SAP model, estimated 

directly from the nominal weight and dimension of all components of the testbed, is shown in 

column (2). By this approach, the total weight of the testbed building was 5122kN (1151 kips). 

However, as mentioned earlier, the measured weight of the testbed was actually 5250 kN (1180 

kips), which was determined from the uplift investigation of the TP bearings. To match the 

measured weight, the weight in the numerical frame model was increased by a factor of 

5220/5122 = 1.019. The weight increase was assumed to be uniform over all nodes of the 

model. Column (3) indicates the factored weight at all floors applied to the OpenSees model. 

Columns (4) and (5) indicate the mass eccentricity at every floor, or distance from the geometric 

center of the building to the center of mass. In general, the center of mass of each floor shifted 

toward the North-East relative to the geometric center. At the base, the center of mass shifted 

toward the West due to the weight of the column bases on the West side as well as the 

staircase on the SouthWest. The Y direction eccentricity at floor 5 increased relative to other 

floors due to the absence of the staircase in the 5th story. The eccentricity was greatest at the 

(a) (b)
147

roof due to the added steel blocks. As mentioned previously, this added weight was excessive 

compared to typical roof mounted equipment and influenced the seismic response of the 

testbed building.

Table 8-1 Weight and eccentricity (distance from geometric center to center of mass) of 

numerical simulation model. 

Floor Weight from Modified weight Eccentricity

SAP (kN) (kN) X(cm) Y(cm)

(1) (2) (3) (4) (5)

Roof 1153 1175 9 -85

5 771 786 20 -40

4 781 796 21 -24

3 782 797 27 -22

2 792 807 22 -24

Base 842 859 0 31

Sum 5122 5220 16 -32

8.4 Modeling Damping

Rayleigh damping (combining mass and stiffness proportional components) was used to 

represent energy dissipation in the testbed building, based on specified damping ratios at two 

different frequencies. The experimental response of the fixed base building to white noise 

excitations was analyzed to find the periods and damping ratios of natural modes of the 

structure (Sasaki et al., 2012), which are listed in Table 8-2 for the first 3 modes in both 

directions. The Rayleigh damping curve passing through damping ratios of 2.2% at periods of 

0.70 sec (frequency of 1.43 Hz), corresponding to the first mode period, and 0.15 sec (6.67 Hz) 

(Figure 8-7), was found to give a good match between experimental and numerical results of the 

fixed-base building.

Past experiments have shown that the damping ratios determined by the method described 

above include damping in the hydraulic actuators of the earthquake simulator and thus overestimate the actual damping in the structural system. If determined from free vibration analysis, 

damping ratios on the order of 1% or less are expected for the first few modes of bare steel 

frame systems (e.g. Uang and Bertero 1986, Whittaker et al. 1990). The damping in the 

hydraulic actuators was not accounted for in the numerical modeling of the system. However, 

additional damping in the testbed building was present compared to a bare steel frame due to 

the concrete floor system, exterior concrete cladding, nonstructural components (partition walls, 

ceilings and piping system) and contents. Thus, the level of observed damping in the testbed 

building was partially justified.

The final “best fit” damping coefficients for the building with hybrid LR isolation system were 

determined by trial and error. The basic strategy behind the selection was to control damping 

between the periods of 0.1 sec (10 Hz) and 2 sec (0.5 Hz) and, which included the major 

response components that were observed in the floor spectra (see Section 9.3). The damping 

ratios were fixed at 1.9% and 1.93% at 0.15 sec (6.66 Hz) and 2.0 sec (0.5 Hz). The Rayleigh 

damping curve for the isolated building configuration is compared to that for the fixed-base 

building configuration in Figure 8-7.
148

Table 8-2 Experimentally determined natural periods and damping ratios of the fixedbase building. 

White noise X White noise Y White noise 3D

Period 

(s)

Damping 

ratio (%)

Period 

(s)

Damping 

ratio (%)

Period 

(s)

Damping 

ratio (%)

Mode 1 X 0.65 3.3 n/a n/a 0.68 4.1

Mode 2 X 0.20 1.6 n/a n/a 0.21 2.0

Mode 3 X 0.11 3.3 n/a n/a 0.11 3.7

Mode 1 Y n/a n/a 0.68 2.5 0.69 3.5

Mode 2 Y n/a n/a 0.21 1.7 0.21 1.9

Mode 3 Y n/a n/a 0.11 2.6 0.11 3.6

Figure 8-7 Rayleigh damping model for the fixed-base and hybrid LR isolated building 

model

The calibrated Rayleigh damping model produced low damping ratios at frequencies around 

3 Hz (Figure 8-7), which were the frequencies of the first structural modes of the isolated 

structure in both directions, so that these frequency components of the numerically simulated 

response tended to be amplified compared to the experimental data. To solve this difficulty, 

additional dampers were added to apply extra damping to these modes (Dao 2012). From 

modal analysis, the relative horizontal displacement between the base and roof in the 1st

structural mode in each direction was observed to be much larger than in other modes. Thus, 

additional dampers were connected between the center of stiffness nodes at the base and roof 

in each direction as shown in Figure 8-8. At these locations, the displacements in the 1st

torsional mode were zero. The damping coefficient cc for the damper in a given direction was 

computed as (Dao 2012):

 cc = 2.ζζnn.ωωnn.mmnnnn

∗

ccnnnn̅ (8.7)

where = mode number of the modified damping mode; ζζ = desired additional damping ratio; 

ωω = the angular frequency of the tth mode; mm∗ = modal mass of the tth mode; and cc̅ = a 

constant dependent on tth mode shape, computed by:

0 2 4 6 8 10

0

2

4

6

8

10

Frequency (Hz)

Damping Ratio (%)

Isolated-base model

Fixed-base model

Fixed-base test
149

 cc̅ = �φφpp − φφqq �

2 (8.8)

where φφpp , φφqq = horizontal displacements at base and roof in the tth mode shape. 

Computation of �cc̅ � for the first 14 modes, including off diagonal terms, demonstrated that the 

extra damping contributed primarily to the intended mode (Dao 2012). The selected value of ζζ

for numerical simulation was 2% for the for the hybrid LR isolation configuration.

Figure 8-8 Additional damper for adjusting the damping coefficient of the 2nd mode in 

the XX direction

cc

XX

ZZ

151

9. COMPARISON OF EXPERIMENTAL AND NUMERICAL SIMULATION

In this chapter, we compare the predicted responses of the hybrid LR isolated building (both 

isolation system and structure) using the OpenSees model described in Chapters 7 and 8 to the 

experimental results. Response quantities examined include the isolator displacement and 

force, story drifts, floor accelerations, and floor response spectra. Four different input excitations 

have been chosen as a representative variety sample. These comparisons are shown for the 

following input excitations: El Centro 130%, Vogtle 100%, Vogtle 175%, Diablo Canyon 95%.

Vogtle was selected because of its great interest to the research sponsor, and two different 

intensities were chosen to approximately represent the DBE and the beyond DBE. Diablo 

Canyon was chosen because of its secondary interest to the sponsor, and because it produced 

the largest displacement demand in the bearings. El Centro was chosen to be representative of 

a smaller earthquake. The earthquake records used as input to the numerical model are the 

recorded output at the base of the earthquake simulator. The target ground excitation for Diablo 

is bidirectional only (XY) excitation, while the other three excitations include vertical input. These 

motions have been selected as a representative variety sample, and focus on the particular 

excitations that are of greatest interest to the research sponsor. Reports comparing the 

experimental and numerical simulation results for every excitation are permanently archived in 

the NEEShub Project Warehouse (Ryan et al. 2012).

9.1 Isolator Response

Demonstration that the isolation system and overall structure demands of safety related nuclear 

structures can be accurately determined in analysis and design is an important milestone 

toward the acceptance and implementation of seismic isolation for nuclear structures. As 

discussed in Chapter 7, replication of the isolation system response using a single bearing 

model with physically determined parameters has been hampered by circumstances unique to 

the test program that would not normally be encountered in practice. These circumstances are: 

1) the bearings were designed with an unusually small lead plug, resulting in pinching of the 

bearing hysteresis loops that is difficult to represent with a single bilinear model. 2) The 

sequence of closely spaced trials caused heating of the lead plugs. Since the bearing 

temperature was not measured, the influence of heating on the bearing response cannot be 

quantified. Under normal field conditions and in a strong earthquake, LR bearings would be 

activated in an unheated state. 3) The experimental simulation results include small intensity 

ground motions that induce limited displacement demands in the isolation systems. A bilinear 

bearing model would not be expected to replicate isolator demands in the small intensity 

simulations well.

As discussed in Chapter 7, in lieu of numerical simulation with a single bearing model, we 

independently calibrate the bearing modeling parameters for each experimental simulation to 

represent the largest displacement cycle that was observed. Use of this technique is helpful to 

interpret the data from this test program, but does not imply that a single bilinear model with 

physical parameters would not be suitable over a wide range of intensities in practice. Also 

discussed in Chapter 7, for all subsequent numerical simulation we substitute a model with Qd

equal to 85% of the best fit characterized model, based on the trial-and-error observation that it 

improves the results.

Displacement traces, displacement histories, and bearing force vs. displacement (hysteresis 

loops) for the four selected input motions are shown in Figures 9-1 to 9-8. This adjusted 

characterized model – with its acknowledged limitations – generally predicts the amplitude of 

the peak vector displacement in each LR bearing quite well; a prediction within 5% of the 

recorded peak is not uncommon. The model also tends to represent the largest cycles in the 

displacement traces well, and capture obvious differences among the 4 LR bearings that are the 
152

result of the torsion in the system. For example, in Vogtle 100%, the numerical simulation 

correctly picks up a linear displacement along a diagonal axis in the East (E) and South (S) 

bearings, but more of a circular orbit pattern in the North (N) and West (W) bearings (Figure 9-

3). Plots of the displacement history best illustrate the accuracy of the model throughout the 

records. As discussed in Chapter 7, the predicted displacement is most accurate at cycles close 

to the peak intensity for which it has been optimized. Thus, the numerical prediction of the 

displacement history is not as accurate over small intensity cycles in large intensity records (e.g. 

Vogtle 175%, Figure 9-5). However, the peak displacement – which is the most critical response 

parameter in the isolation system – is predicted very well by the model, and the prediction in the 

small displacement range is not too important. The numerical simulation also predicts well the 

peak base rotation angle in the isolation system, which is more sensitive to modeling errors. An 

example is shown for Vogtle 175% in Figure 9-9, which compares the history of base rotation 

angle as predicted by numerical and experimental simulation. As discussed earlier, the 

rotational demands in the isolation system that occurred in this testbed structure resulted from 

the limitations on number and placement of LR bearings, and could expect to be eliminated in 

practical buildings with many bearings distributed across the base plan. 

The hysteresis loops confirm that, by inspection, the adjusted characterized numerical model is 

a good fit to the experimental data in most cases, and the fitted value of Kd looks reasonable. 

The numerical model tends to underestimate the force at the peak displacement. (e.g. consider 

the negative excursion of the bearings in both x and y-directions for El Centro, Figure 9-2), 

which results from the pinching behavior of the hysteresis loops due to a small diameter lead 

plug. Because of the pinching behavior, the best fit bilinear model parameter for Qd exceeds the 

observed yield strength Qd near the center of the loops, but is less than the physical Qd

(applicable at large displacements). Also, the numerically simulated loops show a tendency to 

be “wavy” compared to the experimental data. For example, in the El Centro motion, on the 

largest negative excursion in the x-direction, the numerically predicted force dips suddenly on 

the return for all 4 LR bearings, while the actual force recorded during the experimental 

simulation does not (Figure 9-2). In Vogtle 100%, the numerical hysteresis loops (Figure 9-4) 

are characterized by rapid up and down variations of force on a large displacement excursion 

that are absent from the experimental loops. We are not sure, but believe the waviness in the 

numerical hysteresis loops is related to the algorithm for bidirectional interaction. 
153

Figure 9-1 Comparison of numerical and experimental simulation for El Centro 130%; 

displacement trace and displacement histories (x and y) in each LR bearing, 

labeled by position (E, S, N, W). Numerical simulation uses the characterized 

model with 85% of the characterized value for Qd

-20

-10

0

10

20

Displacement Trace

19.3

20.6

E

-15

0

15

-15

0

15

X

Y

E

Displacement History

-20

-10

0

10

20

19.5

S 19.7

-15

0

15

-15

0

15

X

Y

S

-20

-10

0

10

20

18.2

N 17

-X Displacement (cm)

-15

0

15

-15

0

15

X

Y

N

Displacement (cm)

-20 -10 0 10 20

-20

-10

0

10

20

18.5

16.6

W

Y Displacement (cm)

5 10 15 20 25

-15

0

15

-15

0

15

X

Y

W

Time (sec)

Numerical

Experimental
154

Figure 9-2 Comparison of numerical and experimental simulation for El Centro 130%; 

bearing force vs. displacement or hysteresis loops (x and y) in each LR 

bearing, labeled by position (E, S, N, W). Numerical simulation uses the 

characterized model with 85% of the characterized value for Qd

-200

-100

0

100

200

E

Force vs Disp. - X

-200

-100

0

100

200

E

Force vs Disp. - Y

-200

-100

0

100

200

S

-200

-100

0

100

200

S

-200

-100

0

100

200

Force (kN)

N

-200

-100

0

100

200

N

-20 -10 0 10 20

-200

-100

0

100

200

W

Displacement (cm)

-20 -10 0 10 20

-200

-100

0

100

200

W

Displacement (cm)

Numerical

Experimental
155

Figure 9-3 Comparison of numerical and experimental simulation for Vogtle 100%; 

displacement trace and displacement histories (x and y) in each LR bearing, 

labeled by position (E, S, N, W). Numerical simulation uses the characterized 

model with 85% of the characterized value for Qd

-20

-10

0

10

20

Displacement Trace

27.3

26.5

E

-15

0

15

-15

0

15

X

Y

E

Displacement History

-20

-10

0

10

20

21

21.7

S

-15

0

15

-15

0

15

X

Y

S

-20

-10

0

10

20 25.6

25.6

N

-X Displacement (cm)

-15

0

15

-15

0

15

X

Y

N

Displacement (cm)

-20 -10 0 10 20

-20

-10

0

10

20

21

21.7

W

Y Displacement (cm)

5 10 15 20 25

-15

0

15

-15

0

15

X

Y

W

Time (sec)

Numerical

Experimental
156

Figure 9-4 Comparison displacement (x and y) histories of numerical and 

 experimental simulation for Vogtle 100% in each LR bearing labeled my 

position (E, S, N, W).Numerical simulation uses the characterized model with 

85% of the characterized value for Qd

-150

0

150

E

Force vs Disp. - X

-150

0

150

E

Force vs Disp. - Y

-150

0

150

S

-150

0

150

S

-150

0

150

N

Force (kN)

-150

0

150

N

-15 0 15

-150

0

150

W

Displacement (cm)

-15 0 15

-150

0

150

W

Displacement (cm)

Numerical

Experimental
157

\ 

Figure 9-5 Comparison of numerical and experimental simulation for Vogtle 175%; 

displacement trace and displacement histories (x and y) in each LR bearing, 

labeled by position (E, S, N, W). Numerical simulation uses the characterized 

model with 85% of the characterized value for Qd

-40

-20

0

20

40

Displacement Trace

48.1

50.5

E -30

0

30

-30

0

30

X

Y

E

Displacement History

-40

-20

0

20

40 46.4

44.7

S -30

0

30

-30

0

30

X

Y

S

-40

-20

0

20

40 47.6 49.4

N

-X Displacement (cm)

-30

0

30

-30

0

30

X

Y

N

Displacement (cm)

-40 -20 0 20 40

-40

-20

0

20

40 46.4

44.6

W

Y Displacement (cm)

5 10 15 20 25

-30

0

30

-30

0

30

X

Y

W

Time (sec)

Numerical

Experimental
158

\ 

Figure 9-6 Comparison of numerical and experimental simulation for Vogtle 175%; 

displacement trace and displacement histories (x and y) in each LR bearing, 

labeled by position (E, S, N, W). Numerical simulation uses the characterized 

model with 85% of the characterized value for Qd

-400

-200

0

200

400 E

Force vs Disp. - X

-400

-200

0

200

400 E

Force vs Disp. - Y

-400

-200

0

200

400 S

-400

-200

0

200

400 S

-400

-200

0

200

400 N

Force (kN)

-400

-200

0

200

400 N

-40 -20 0 20 40

-400

-200

0

200

400 W

Displacement (cm)

-40 -20 0 20 40

-400

-200

0

200

400 W

Displacement (cm)

Numerical

Experimental
159

Figure 9-7 Comparison of numerical and experimental simulation for Diablo 95%; 

displacement trace and displacement histories (x and y) in each LR bearing, 

labeled by position (E, S, N, W). Numerical simulation uses the characterized 

model with 85% of the characterized value for Qd

-50

-25

0

25

50

Displacement Trace

54.8

54.7

E -30

0

30

-30

0

30

X

Y

E

Displacement History

-50

-25

0

25

50 49.6

51.1

S -30

0

30

-30

0

30

X

Y

S

-50

-25

0

25

50 45

43.7

N

-X Displacement (cm)

-30

0

30

-30

0

30

X

Y

N

Displacement (cm)

-50 -25 0 25 50

-50

-25

0

25

50 41.9

41.2

W

Y Displacement (cm)

10 15 20 25

-30

0

30

-30

0

30

X

Y

W

Time (sec)

Numerical

Experimental
160

Figure 9-8 Comparison displacement (x and y) histories of numerical and 

 experimental simulation for Diablo 95% in each LR bearing labeled my 

position (E, S, N, W).in each LR bearing, labeled by position (E, S, N, W). 

Numerical simulation uses the characterized model with 85% of the 

characterized value for Qd

-400

-200

0

200

400 E

Force vs Disp. - X

-400

-200

0

200

400 E

Force vs Disp. - Y

-400

-200

0

200

400 S

-400

-200

0

200

400 S

-400

-200

0

200

400 N

Force (kN)

-400

-200

0

200

400 N

-60 -40 -20 0 20 40 60

-400

-200

0

200

400 W

Displacement (cm)

-60 -40 -20 0 20 40 60

-400

-200

0

200

400 W

Displacement (cm)

Numerical

Experimental
161

Figure 9-9 Comparison of numerical and experimental simulation for Vogtle 175%; 

history of base rotation angle. Analysis uses the characterized model with 

85% of the characterized value for Qd

There is no evidence of loss of axial load carrying capacity in the bearing hysteresis loops. Such 

evidence would include a decrease of the stiffness at large horizontal displacements that 

occurred in combination with high overturning induced axial loads, and the inability of the 

numerical model, which does not include the interaction of horizontal shear and axial force, to 

simulate the response. Neither of these effects are observed in any of the hysteresis loops. In 

summary, the overall ability of the characterized model to predict the bearing displacement and 

bearing force is sufficient that this model can be expected to provide insight into the structural 

response, which is examined next.

Simulations of the axial force in the isolators as a function of time are not shown. Attempts to 

replicate the bearing axial force variation with the present numerical simulation model were 

unsuccessful. It may be possible to replicate the axial force variation and load transfer with a 

bearing model that predicts the horizontal-vertical coupling in the bearing. Such a model could 

also predict axial force instability in the bearing. Such models exist (e.g. Ryan et al. 2005, 

Yamamoto et al. 2009) but were not readily available to the authors and could not be 

implemented within the time constraints of the report. Because of the difficulties in reliable 

prediction of the LR bearing axial forces when used as a hybrid system, we recommend that 

hybrid LR isolation systems need not be applied to nuclear facilities. Isolation systems 

consisting solely of LR bearings and designed outside of the constraints of this test program will 

have similar response characteristics without the load transfer effects.

9.2 Modal Analysis

An analysis of the mode shapes and frequencies of the numerical model is useful for 

interpreting the floor spectra, and understanding the response to combined horizontal and 

vertical excitations (explored in depth in Chapter 10). For modal analysis, we assign equivalent 

linear parameters to the isolator models, which are horizontal stiffness k = 8.46 kN/cm (4.83 

kip/in) for the LR bearings (corresponding to period T = 2.5 sec), horizontal stiffness k = 0.024 

kN/cm (0.014 kip/in) for the CL bearings (equal to the initial stiffness, Section 7.2.1), and vertical 

5 10 15 20 25

-0.02

-0.01

0

0.01

0.02

Time (sec)

Rotation (rad)

0.019

0.019

Numerical

Experimental
162

stiffness equal to the calibrated model compression stiffness for both LR bearings (Section 

7.1.2) and CL bearings (Section 7.2.2). Because the isolator models are nonlinear, the actual 

mode shapes and frequencies of the testbed building vary dynamically such that the modes 

shown here are at best approximate. However, the higher modes, which are of greatest interest 

in the subsequent discussion, are relatively insensitive to variation of the isolator stiffness over a 

reasonable working range, such that an analysis based on equivalent linear mode shapes is 

meaningful. We will demonstrate that some of the higher frequency modes are observed in the 

responses of the building in some input excitations. 

Figures 9-10 to 9-14 illustrate the first 20 modes of the numerical model, along with their 

associated frequencies and periods. For each mode, the following views are plotted: the 

deflected shape of the testbed framing in 3-dimensional space, the deflected shape of each of 3 

frames in the x-direction, and the deflected shape of each of 3 frames in the y-direction. The 

undeflected shape of the system is shown in dashed lines for reference. The isolator 

displacement (displacement at base), frame deflection and panel zone deflection can be seen in 

the plots. 

The first 3 modes are isolation modes that induce mainly horizontal deflection in the isolators, 

and almost no deflection in the testbed framing. Mode 1 (T = 2.72 sec) is primarily an x-direction 

mode, mode 2 (T = 2.60 sec) is primarily a y-direction mode, and mode 3 (T = 2.26 sec) is a 

torsional mode. Mode 1 also includes torsional deflection due to several sources of stiffness and 

mass irregularity in the y-direction (e.g. unequal bay widths in y-direction, staircase at the SW 

corner of the building, and supplementary mass at the roof level on the E half of the building). 

The rotational demands observed at the base (Section 5.2) are not unexpected for this test 

structure, since the isolation system could not be configured to resist torsion, as pointed out in 

Section 3.3.2.

Modes 4-6 are the first structural modes, characterized by horizontal deflection in the isolation 

system, and linear deflection across the structural system with a single node (point of zero 

deformation) at approximately the 4th floor. Based on the assumed equivalent linear stiffness of 

the isolators, mode 4 (T = 0.38 sec) is a y-direction mode, mode 5 (T = 0.36 sec) is an xdirection mode, and mode 6 (T = 0.29 sec) is a torsional mode. Likewise, modes 7, 8 and 10 are 

the second structural modes, characterized by structural deflection with 2 nodes at 

approximately the 2nd and 5th floors. Analogous to previous sets, modes 7 (T = 0.18 sec) and 8 

(T = 0.17 sec) are primarily y-direction and x-direction modes, respectively, while mode 10 (T = 

0.14 sec) is a torsional mode. Modes 13 (T = 0.104 sec) and 14 (T = 0.100 sec) comprise a 3rd

set of structural modes, with 3 nodes each in the 1st story, 3rd story and 5th story. Although 

these are primarily lateral modes, they also contain some notable slab vibration at the 4th and 

roof levels. A corresponding torsional mode does not appear in the first 20 modes that have 

been plotted.
163

Figure 9-10 Modes 1-4 deflected shape of all the frames in 3-dimensional, x- and ydirection, along with their associated frequencies and periods. The 

undeflected shape of the system is shown in dashed lines

 X Y

 X Y

 X Y

 

 X Y
164

 

Figure 9-11 Modes 5-8 deflected shape of all the frames in 3-dimensional, x- and ydirection, along with their associated frequencies and periods. The 

undeflected shape of the system is shown in dashed lines

 X Y

 X Y

 X Y

 

 X Y
165

Figure 9-12 Modes 9-12 deflected shape of all the frames in 3-dimensional, x- and ydirection, along with their associated frequencies and periods. The 

undeflected shape of the system is shown in dashed lines.

 X Y

 X Y

 X Y

 

 X Y
166

Figure 9-13 Modes 13-16 deflected shape of all the frames in 3-dimensional, x- and ydirection, along with their associated frequencies and periods. The 

undeflected shape of the system is shown in dashed lines

 X Y

 X Y

 X Y 

 X Y
167

Figure 9-14 Modes 17-20 deflected shape of all the frames in 3-dimensional, x- and ydirection, along with their associated frequencies and periods. The 

undeflected shape of the system is shown in dashed lines

 X Y

 X Y

 X Y

 

 X Y
168

As observed in Figures 9-12 to 9-14, the remaining modes (9,11,12,15-20) are local modes 

dominated by vertical vibration of individual slabs. Such modes are usually ignored when 

evaluating the dynamic characteristics of a building, but become more significant when the 

response to vertical excitation is evaluated. If movement of the horizontal degrees of freedom 

(DOFs) is present in the local modes, the horizontal floor accelerations and floor spectra will be 

amplified when the system is subjected to vertical excitation compared to horizontal only 

excitation. This phenomenon has not often been encountered in the literature. Mode 9 (T = 0.14 

sec, Figure 9-12) could be interpreted as the fundamental vertical mode. Uncharacteristic of a 

typical first vertical mode, Mode 9 is characterized primarily by vertical slab vibration at the roof 

level at the location directly beneath the supplementary mass. The supplementary mass has 

been confirmed (from SAP models not presented here) to increase the mass participation of 

Mode 9 as well as the slab vibration at the roof level relative to a typical building; however, the 

movement of horizontal DOFs in Mode 9 appears to be negligible (Figure 9-12). The influence 

of the supplementary mass on the response of the isolated testbed building is explored 

numerically in Chapter 10. Mode 11 (T = 0.107 sec) is a local vertical mode dominated by slab 

vibration at the 5th floor. Mode 12 (T = 0.105 sec) is a local mode dominated by slab vibration at 

the roof level that also contains movement of horizontal DOFs, especially at the 2nd, 3rd and 5th

floors (Figure 9-12). Modes 15-20 (Figures 9-13 and 9-14) all appear to be local modes 

characterized by vertical slab vibration at several levels, but with little to no movement of 

horizontal DOFs except in Mode 15 (T = 0.098 sec) in the y-direction. Modes 15 and above 

have low mass participation.

9.3 Structural Response

In this section, we compare the structural responses predicted by the numerical model to those 

recorded from the sensors. The x and y-direction roof acceleration histories, peak acceleration 

profiles (peak acceleration vs. story level), 2nd story drift histories, and peak drift profiles (peak 

drift vs. story level) are plotted for the four selected input motions in Figures 9-15 to 9-18. The 

response quantities are reported at the geometric center of the structure, obtained for the 

experimental data by averaging or linear interpolation of the filtered data from multiple sensors. 

The roof and 2nd story are chosen for plotting the acceleration and drift histories because the 

largest demands are observed at these locations, respectively. 

Both low frequency and higher frequency components vibrations are visible in the roof 

acceleration and 2nd story drift histories (Figures 9-15 to 9-18). These signals have a low 

frequency motion vibration appears as a slow moving wave at the isolation frequency that 

determines the amplitude of oscillation through them that is indicative of the isolation frequency, 

while the higher frequency motions appear as rapid oscillations about the slower moving wave. 

As expected, the low frequency motion is strong in the drift histories, but obscured by higher 

frequency components in the acceleration histories. The observed isolation frequency is lower 

for El Centro (Figure 9-15) than for the other records, to reflect the nonlinearity of the system 

and thus the change in frequency with amplitude. A second component, strong in all signals, is 

deduced by visual inspection to be around 3 Hz (0.33 sec), and thus represents the first 

structural mode.
169

Figure 9-15 Comparison of numerical and experimental simulation for El Centro 130%; x 

and y-direction roof acceleration histories, peak acceleration profiles, 2nd 

story drift histories, and peak drift profiles. Numerical simulation uses the 

characterized model with 85% of the characterized value for Qd

5 10 15 20 25

-0.2

-0.1

0

0.1

0.2

Acc. 6th Floor - X

Acceleration (g)

Time (sec)

5 10 15 20 25

-0.2

-0.1

0

0.1

0.2

Acc. 6th Floor - Y

Time (sec)

0 0.2 0.4 0.6

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - X

Peak Acceleration (g)

0 0.2 0.4 0.6

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - Y

Peak Acceleration (g)

5 10 15 20 25

-0.2

-0.1

0

0.1

0.2

Story Drift 2nd Floor - X

Time (sec)

Drift (%)

5 10 15 20 25

-0.2

-0.1

0

0.1

0.2

Story Drift 2nd Floor - Y

Time (sec)

0 0.1 0.2 0.3 0.4

Base

1st

2nd

3rd

4th

5th

Story Drift Profile - X

Drift (%)

0 0.1 0.2 0.3 0.4

Base

1st

2nd

3rd

4th

5th

Story Drift Profile - Y

Drift (%)

Numerical

Experimental
170

 

Figure 9-16 Comparison of numerical and experimental simulation for Vogtle 100%; x 

and y-direction roof acceleration histories, peak acceleration profiles, 2nd 

story drift histories, and peak drift profiles. Numerical simulation uses the 

characterized model with 85% of the characterized value for Qd

5 10 15 20 25

-0.4

-0.2

0

0.2

0.4

Acc. 6th Floor - X

Acceleration (g)

Time (sec)

5 10 15 20 25

-0.4

-0.2

0

0.2

0.4

Acc. 6th Floor - Y

Time (sec)

0 0.2 0.4 0.6

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - X

Peak Acceleration (g)

0 0.2 0.4 0.6

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - Y

Peak Acceleration (g)

5 10 15 20 25

-0.2

-0.1

0

0.1

0.2

Story Drift 2nd Floor - X

Time (sec)

Drift (%)

5 10 15 20 25

-0.2

-0.1

0

0.1

0.2

Story Drift 2nd Floor - Y

Time (sec)

0 0.1 0.2 0.3 0.4

Base

1st

2nd

3rd

4th

5th

Story Drift Profile - X

Drift (%)

0 0.1 0.2 0.3 0.4

Base

1st

2nd

3rd

4th

5th

Story Drift Profile - Y

Drift (%)

Numerical

Experimental
171

 

Figure 9-17 Comparison of numerical and experimental simulation for Vogtle 175%; x 

and y-direction roof acceleration histories, peak acceleration profiles, 2nd 

story drift histories, and peak drift profiles. Numerical simulation uses the 

characterized model with 85% of the characterized value for Qd

5 10 15 20 25

-0.5

0

0.5

Acc. 6th Floor - X

Acceleration (g)

Time (sec)

5 10 15 20 25

-0.5

0

0.5

Acc. 6th Floor - Y

Time (sec)

0 0.5 1

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - X

Peak Acceleration (g)

0 0.5 1

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - Y

Peak Acceleration (g)

5 10 15 20 25

-0.2

-0.1

0

0.1

0.2

Story Drift 2nd Floor - X

Time (sec)

Drift (%)

5 10 15 20 25

-0.2

-0.1

0

0.1

0.2

Story Drift 2nd Floor - Y

Time (sec)

0 0.1 0.2 0.3 0.4

Base

1st

2nd

3rd

4th

5th

Story Drift Profile - X

Drift (%)

0 0.1 0.2 0.3 0.4

Base

1st

2nd

3rd

4th

5th

Story Drift Profile - Y

Drift (%)

Numerical

Experimental
172

 

Figure 9-18 Comparison of numerical and experimental simulation for Diablo 95%; x and 

y-direction roof acceleration histories, peak acceleration profiles, 2nd story 

drift histories, and peak drift profiles. Numerical simulation uses the 

characterized model with 85% of the characterized value for Qd

10 15 20 25

-0.5

0

0.5

Acc. 6th Floor - X

Acceleration (g)

Time (sec)

10 15 20 25

-0.5

0

0.5

Acc. 6th Floor - Y

Time (sec)

0 0.5 1

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - X

Peak Acceleration (g)

0 0.5 1

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - Y

Peak Acceleration (g)

10 15 20 25

-0.2

-0.1

0

0.1

0.2

Story Drift 2nd Floor - X

Time (sec)

Drift (%)

10 15 20 25

-0.2

-0.1

0

0.1

0.2

Story Drift 2nd Floor - Y

Time (sec)

0 0.1 0.2 0.3 0.4

Base

1st

2nd

3rd

4th

5th

Story Drift Profile - X

Drift (%)

0 0.1 0.2 0.3 0.4

Base

1st

2nd

3rd

4th

5th

Story Drift Profile - Y

Drift (%)

Numerical

Experimental
173

The frequency components in the experimental response histories are clearly replicated in the 

numerically simulated responses, which is strong validation of the accuracy of the numerical 

simulation. The peak amplitude of response histories are difficult to simulate numerically due to 

inevitable noise and spikes in recorded experimental data, as well as the sensitivity of the 

response to the low values of damping in the calibrated numerical model. Given these 

difficulties, the numerically simulated peak amplitudes of floor acceleration and story drift are 

quite accurate, generally within about 20% of the experimentally observed values.

As further indication of the effectiveness of the numerical simulation, the numerically simulated 

acceleration profiles generally follow the trends observed in the experimental acceleration 

profiles. For instance, for all of the records except Diablo 95% (Figure 9-18), the accelerations in 

the y-direction are larger than in the x-direction, replicated in both experimental and numerical 

simulation data. The acceleration profile pattern in the x-direction is characterized by maxima at 

the base and the roof and a minimum at the 4th floor. This pattern is clearly indicative of the 

combination of an isolation mode (uniform) and a first structural mode (linear with a node at the 

4th floor). A larger slope in the acceleration profile (or greater difference between the maximum 

and minimum acceleration over the height) indicates greater participation of the first structural 

mode. The shape of the acceleration profile through these control points is more jagged in the 

lower intensity motions (El Centro 130% and Vogtle 100%, Figures 9-15 and 9-16) and 

smoother in the larger motions (Vogtle 175% and Diablo 95%, Figures 9-17 and 9-18). All these 

trends are well captured by numerical simulation. In the y-direction, the acceleration profile 

shapes from the experimental data are not as consistent from record to record, which suggests 

that other modes may be participating. The numerical simulation seems to pick up a false peak 

at the 5th floor in the y-direction for some records such as Vogtle 100% (Figure 9-16). For further 

insight, the 5% damped x and y-direction floor spectra, obtained from the numerical and 

experimentally obtained floor acceleration histories, are plotted for every floor in Figures 9-19 to 

9-22 for the same four selected input motions. Because the target input acceleration for Diablo 

95% is bidirectional (without vertical), it is easiest to analyze. Peaks occurring at periods below 

about 0.1 sec are ignored and high frequency components (above 50 Hz or below 0.02 sec) are 

filtered out of the signals. In the Diablo 95% floor spectra (Figure 9-22), three distinct peaks 

appear. A first peak appears on every floor at a period between 2 and 3 sec (0.33 to 0.5 Hz), 

and represents the isolation mode in each direction. As mentioned previously, the period of this 

peak varies from record to record since the isolation system is nonlinear. A second peak 

appears on all floors except the 4th at a period between 0.3 and 0.4 sec (2.5 to 3.3 Hz), and 

represents the first structural mode in each direction. These modes (4 and 5), plotted in Figure 

9-10 and 9-11 each have a node at the 4th floor. A third peak appears on all floors except the 2nd

and 5th just below a period of 0.2 sec (5 Hz), and represents the second structural mode in each 

direction. These modes (7 and 8), plotted in Figure 9-11, each have nodes at the 2nd and 5th

floor. These floor spectra confirm that participation of a 2nd structural mode in each direction, 

which could not be detected by visual inspection in the roof acceleration histories, is nonnegligible in the acceleration response. 
174

Figure 9-19 Comparison of numerical and experimental simulation for El Centro 130%; x 

and y-direction spectral accelerations for 1st through 6th floors. Numerical 

simulation uses the characterized model with 85% of the characterized value 

for Qd

0

0.5

1

F1 - X F1 - Y

0

0.5

1

F2 - X F2 - Y

0

0.5

1

F3 - X F3 - Y

0

0.5

1

F4 - X F4 - Y

0

0.5

1

F5 - X F5 - Y

10-3 10-2 10-1 100101 0

0.5

1

F6 - X

Period (s)

10-3 10-2 10-1 100101

F6 - Y

Period (s)

10-3 10-2 10-1 100101

F6 - Y

Period (s)

10-3 10-2 10-1 100101 0

0.5

1

F6 - X

Period (s)

F5 - Y

0

0.5

1

F5 - X

F4 - Y

0

0.5

1

F4 - X

Spectral Acceleration (g)

F3 - Y

0

0.5

1

F3 - X

F2 - Y

0

0.5

1

F2 - X

F1 - Y

0

0.5

1

F1 - X

Numerical

Experimental
175

 

Figure 9-20 Comparison of numerical and experimental simulation for Vogtle 100%; x 

and y-direction spectral accelerations for 1st through 6th floors. Numerical 

simulation uses the characterized model with 85% of the characterized value 

for Qd. 

0

0.5

1

F1 - X F1 - Y

0

0.5

1

F2 - X F2 - Y

0

0.5

1

F3 - X F3 - Y

0

0.5

1

F4 - X F4 - Y

0

0.5

1

F5 - X F5 - Y

10-3 10-2 10-1 100101 0

0.5

1

F6 - X

Period (s)

10-3 10-2 10-1 100101

F6 - Y

Period (s)

10-3 10-2 10-1 100101

F6 - Y

Period (s)

10-3 10-2 10-1 100101 0

0.5

1

F6 - X

Period (s)

F5 - Y

0

0.5

1

F5 - X

F4 - Y

0

0.5

1

F4 - X

Spectral Acceleration (g)

F3 - Y

0

0.5

1

F3 - X

F2 - Y

0

0.5

1

F2 - X

F1 - Y

0

0.5

1

F1 - X

Numerical

Experimental
176

 

Figure 9-21 Comparison of numerical and experimental simulation for Vogtle 175%; x 

and y-direction spectral accelerations for 1st through 6th floors. Numerical 

simulation uses the characterized model with 85% of the characterized value 

for Qd

0

1

2

F1 - X F1 - Y

0

1

2

F2 - X F2 - Y

0

1

2

F3 - X F3 - Y

0

1

2

F4 - X F4 - Y

0

1

2

F5 - X F5 - Y

10-3 10-2 10-1 100101 0

1

2

F6 - X

Period (s)

10-3 10-2 10-1 100101

F6 - Y

Period (s)

10-3 10-2 10-1 100101

F6 - Y

Period (s)

10-3 10-2 10-1 100101 0

1

2

F6 - X

Period (s)

F5 - Y

0

1

2

F5 - X

F4 - Y

0

1

2

F4 - X

Spectral Acceleration (g)

F3 - Y

0

1

2

F3 - X

F2 - Y

0

1

2

F2 - X

F1 - Y

0

1

2

F1 - X

Numerical

Experimental
177

 

Figure 9-22 Comparison of numerical and experimental simulation for Diablo 95% (XY); 

x and y-direction spectral accelerations for 1st through 6th floors. Numerical 

simulation uses the characterized model with 85% of the characterized value 

for Qd

0

1

2

F1 - X F1 - Y

0

1

2

F2 - X F2 - Y

0

1

2

F3 - X F3 - Y

0

1

2

F4 - X F4 - Y

0

1

2

F5 - X F5 - Y

10-3 10-2 10-1 100101 0

1

2

F6 - X

Period (s)

10-3 10-2 10-1 100101

F6 - Y

Period (s)

10-3 10-2 10-1 100101

F6 - Y

Period (s)

10-3 10-2 10-1 100101 0

1

2

F6 - X

Period (s)

F5 - Y

0

1

2

F5 - X

F4 - Y

0

1

2

F4 - X

Spectral Acceleration (g)

F3 - Y

0

1

2

F3 - X

F2 - Y

0

1

2

F2 - X

F1 - Y

0

1

2

F1 - X

Numerical

Experimental
178

The same peaks in the floor spectra appear on the same floors at essentially the same periods 

for the three remaining motions (Figures 9-19 to 9-21). In the y-direction, an additional peak 

appears at a period of about 0.1 sec (10 Hz) on some floors. This peak is strongest on the 2nd, 

3rd, 5th and roof floors. The most likely candidate for this peak is Mode 13 (Figure 9-13), which is 

the third structural mode in the y-direction. Although not suppressed on any floor, the horizontal 

deformation of Mode 13 is largest on the 2nd, 3rd, and 5th floors, corresponding approximately to 

the largest peaks in the spectra. Unlike the other structural modes, Mode 13 includes significant 

slab vibration at the 4th floor and (especially) roof, suggesting that this mode is driven by vertical 

excitation. The dominant slab vibration at the roof level is increased by the supplementary mass 

installed at the roof, which is explored further explored through numerical simulation in Chapter 

10. The corresponding x-direction mode (Mode 14) contains much lower levels of slab vibration, 

and thus is not excited to the same extent by the vertical input. Although, small peaks in the xdirection floor spectra can sometimes be observed at this frequency (e.g. Vogtle 175%, Figure 

9-21).

The peaks of the numerically predicted spectra generally occur at the same periods (or very 

close) to the spectra for the experimental data. As expected due to the nonlinearity of the 

isolation system, the periods at which these peaks occur do not exactly match the period of the 

modes shown in Figures 9-10 to 9-14. For instance, the peaks in the spectra occur at periods 

around 0.164 sec (6.1 Hz), while the periods of modes 7 and 8 are predicted to be 0.179 sec 

(5.6 Hz) and 0.169 sec (5.9 Hz), respectively. Since the stiffness assumed for the isolators does 

not reflect the actual frequency excited in a given motion, which changes continuously, the 

discrepancy in the periods of the numerically simulated spectral peaks compared to the modal 

peaks is quite reasonable. Indeed, the first mode period from modal analysis was 2.72 sec, 

while the actual first peaks of the floor spectra are closer to 2 sec for most motions, which 

confirms that a corresponding shift of periods throughout the modes is valid.

Strong higher mode participation is observed in the floor accelerations and floor spectra, which 

is a consequence of the low amount of energy dissipation in the steel frame. As described in 

Section 8.6, the Rayleigh damping model has been calibrated for 1.9% at the representative 

higher mode frequency of 6.6 Hz (0.15 sec). Measured damping ratios on the order of 1-2% are 

reasonable for a steel frame system that remains linear. 

Calibration of the Rayleigh damping coefficients is difficult, which affects the accuracy of the 

numerical model in predicting the spectral peaks. The damping has been calibrated to achieve 

the best fit (by inspection) over all 15 simulations, but only a few are shown here. In the 

simulations shown here, the numerical model tends to be overdamped for the first structural 

mode (it underestimates the peak) and is underdamped for the second and higher structural 

modes (it overestimates the peaks). Recall that an additional damper was included in the 

numerical model to represent the first structural mode (Section 8.6). The calibrated value of this 

damper leads to accurate estimates of the first structural mode peaks for Vogtle 100% (Figure

9-20) and El Centro 130% x-direction (Figure 9-19), but underestimates the peaks for the other 

motions. Likewise, the calibrated Rayleigh damping coefficient at higher frequencies (2nd and 3rd

structural modes) leads to accurate or over-estimated spectral peaks for El Centro 130% and 

both Vogtle records in the y-direction, but under-estimated spectral peaks for the rest. Thus, the 

observed energy dissipation in the system – quantified in terms of a viscous damping model – 

seems to vary from trial to trial, and no single assumption for the damping model is perfect. 

Improved accuracy of numerical simulation results relative to the experimental data may be 

achievable with computer programs that construct a damping matrix (even for nonlinear 

analysis) based on independent damping ratios in individual modes. Such approaches are 

infeasible for design since the damping properties of the building are not known in such detail.
179

Overall, the numerical simulation matches the experimental data with sufficient accuracy to a) 

build confidence in the modeling techniques used by the profession, and b) provide insight to 

help interpret the response of the test structure, as has been described above.

As discussed in Chapter 2, a nuclear facility would be substantially stiffer than the tested steel 

moment frame structure, with a natural period in the range of 0.1 to 0.25 sec, while the tested 

structure natural period exceeded 0.6 sec. Relative to the experimental data shown here, a 

stiffer structure with the same isolation period would experience lower overall accelerations and 

reduced higher mode effects. The expected evidence of reduced higher mode effects is a) 

reduced high frequency vibration in the acceleration histories (Figures 9-15 to 9-18), b) 

acceleration profiles that are more uniform (i.e. identical accelerations at all levels from base to 

roof (Figures 9-15 to 9-18), and c) reduction in the peaks of the floor spectra for periods less 

than 1.0 sec. In addition, higher mode effects may be reduced as a result of a simpler (single 

story) structural configuration. Without simulation, the extent of acceleration reduction cannot be 

predicted, but it is expected to be significant. These general trends can be verified by 

fundamental textbooks on the theory of seismic isolation (e.g. Kelly 1997, Chopra 2012).

181

10. INFLUENCE OF VERTICAL EXCITATION ON THE STRUCTURAL 

RESPONSE 

In this chapter, the influence of vertical excitation on the overall response of the isolation system 

and the structure is evaluated. Several XY (horizontal only) and 3D excitations were included in 

the test program, and results for both types of motions have been discussed to date. To provide 

a direct assessment of the influence of vertical excitation, comparative simulations were 

conducted with input motions that included and neglected vertical shaking input. Specifically, the 

Northridge-Rinaldi record – scaled to 88% - was applied to the system as both an XY excitation 

(with target vertical excitation of zero) and as a 3D excitation. The Rinaldi 88% (XY) and Rinaldi 

88% (understood to be 3D) simulations were applied back-to-back in the test program (Table 4-

4). Similarly, the Diablo Canyon record was applied as an XY excitation with 95% scale factor 

(Diablo 95% XY) and as 3D excitation with 80% scale factor (Diablo 80%). The isolator and 

structural responses under the Rinaldi 88% (XY)/Rinaldi 88% and Diablo 95% (XY)/Diablo 80% 

pairings are compared in this chapter.

The substantial supplementary weight mounted at the roof level of the testbed structure, as 

described in Section 2.3, is believed to have influenced the response of the structure due to 

vertical excitations and the subsequent horizontal-vertical coupled response of the structure. As 

such, additional numerical simulations have been conducted using an augmented numerical 

simulation model of the structure. In the augmented model, all supplementary weight at the roof 

level has been removed and relocated at the base level (just above the isolators) in a regular 

configuration, or proportional to the existing mass at the base level. Thus, the augmented model 

essentially retains the same properties of the isolation system, and removes the substantial 

eccentricity and vertical irregularity associated with the roof weight. Numerical simulation results 

using the augmented model are included to provide the perspective on this effect. The 

simulations suggest that a good portion of the horizontal-vertical modal coupling and 

amplification of horizontal floor accelerations resulted from the supplementary mass that was 

placed at the roof level. 

10.1 Response of LR Bearings under Axial Force Variation

The influence of the vertical excitation on the isolation system response is examined first. There 

was no evidence in any of the hysteretic data examined to date (see Section 9.1) that the 

bearing hysteresis loops were affected by loss of axial load carrying capacity under the 

combination of large displacements and increased axial loading due to overturning forces. Since 

the bearings were not subjected to the usual increase in axial force at large lateral 

displacements due to overturning, but rather tended to unload axially at large displacements 

(Section 6.2), loss of lateral load carrying capacity was not expected.

Comparative data from the Rinaldi 88% (XY) and Rinaldi 88% trials suggests that rapid strong 

axial force variation in the bearings affected the horizontal shear force in the bearings in the 

Rinaldi simulation. Figures 10-1 and 10-2 compare the bearing hysteresis loops in X and Ydirections for Rinaldi 88% XY and 3D for a single large displacement cycle. The vertical peak 

ground acceleration (V-PGA) for the 3D excitation is listed. While the displacements in the 

bearings are approximately the same for XY and 3D simulation, a small high frequency 

oscillation was present in the forces for the 3D simulation over certain parts of the hysteresis 

loop that was absent in the XY simulation. This oscillation is distinct from the effects of bolt slip 

documented in Section 6.1. For these simulations, the history of the shear force variation in 

LRB-S is plotted along with axial force variation in Figure 10-3, where the reference cycle shown 

in the hysteresis loops of Figures 10-1 and 10-2 is indicated in bold. Figure 10-3 indicates that 

the high frequency component of the shear force variation was in phase with the axial force 

variation. We do not know the cause of this small shear force variation; it could be a result of 
182

instantaneous variation in confinement of the lead core as the axial force on the bearing cycles 

between a state of tension (axial force > 0 in Figure 10-3) to more than twice the initial static 

load in compression. 

Figure 10-1 Bearing force vs. displacement (hysteresis loops) in x-direction for a single 

large displacement cycle compared for Rinaldi 88% (XY) and Rinaldi 88% (VPGA = 1.21g) excitations

Figure 10-2 Bearing force vs. displacement (hysteresis loops) in y-direction for a single 

large displacement cycle compared for Rinaldi 88% (XY) and Rinaldi 88% (VPGA = 1.21g) excitations
183

Figure 10-3 History of shear force and axial force variation in LRB-S compared for 

Rinaldi 88% (XY) and Rinaldi 88% (V-PGA = 1.21g) excitations

The high frequency shear force variation was not evident in any other records. Data analogous 

to Figures 10-1 to 10-3 (single loop bearing hysteresis and force histories for a single LRB in x, 

y and z-directions) are presented for Diablo 80%/Diablo 95% (XY) in Figures 10-4 to 10-6. 

Because the scale factors for the motions were different, the hysteresis loops do not align 

precisely and the same one-to-one comparison of XY versus 3D excitation was not possible for 

this simulation pairing. In the hysteresis loops of Figures 10-4 and 10-5, some oscillations in the 

shear force were present for both XY and 3D simulations (resulting from bolt slip or bidirectional 

interaction), but no local variations of the shear force could reasonably be attributed to vertical 

excitation.

The peak vertical ground acceleration observed in the Rinaldi 88% test (V-PGA = 1.26g) was 

more than twice that of any other record in the test program (Table 4-7), and local peaks 

exceeding about 0.75g were reached several times during the vertical history, including the part 

of the record when the horizontal acceleration became strong (Figure 4-18). The axial forces in 

the isolators cycled between a state of tension and more than twice the static load in 

compression, which is unusually large. The slight variation of the shear force that is correlated 

with this large intensity and high frequency variation of the axial force on the bearings, 

regardless of its cause, does not appear to be significant.
184

Figure 10-4 Bearing force vs. displacement (hysteresis loops) in x-direction for a single 

cycle compared for Diablo 80% (V-PGA = 0.44g) and Diablo 95% (XY) 

excitations

Figure 10-5 Bearing force vs. displacement (hysteresis loops) in y-direction for a single 

cycle compared for Diablo 80% (V-PGA = 0.44g) and Diablo 95% (XY) 

excitations
185

Figure 10-6 History of shear force and axial force variation in LRB-N compared for 

Diablo 80% (V-PGA = 0.44g) and Diablo 95% (XY) excitations

10.2 Horizontal and Vertical Floor System Vibration

Vertical acceleration produced by the earthquake simulator generated significant vertical 

vibration of the floor system (floor slabs). Figure 10-7 plots vertical accelerations recorded in the 

simulator platform and the middle of the floor slab at the 5th and roof levels for three different 

records: Vogtle 100% (V-PGA = 0.28g), Diablo 80% (V-PGA = 0.44g) and Rinaldi 88% (V-PGA 

= 1.21g). For these records, the peak slab accelerations were amplified relative to the V-PGA, 

and the amplification factor increased slightly as the intensity of V-PGA was increased. The 5th

floor slab acceleration was amplified by a factor of 3.5 to 4.4, while the roof slab acceleration 

was amplified by a factor of 5.4 to 6.3. These amplification factors were determined by the slab 

vibration frequencies and were large due to the relatively low levels of damping. From Figure 

10-7, the slab accelerations appeared to be dominated by single frequency vibration. The 

dominant slab vibration frequencies, confirmed by transfer function analysis (not shown here), 

were about 10 Hz for the 5th floor slab and 7 Hz for the roof slab. Thus, amplification was higher 

at the roof slab than for the 5th floor slab due to its lower natural frequency, which was a result of 

the supplementary weight that was attached at the roof level. 

This suggests that individual local vertical modes were activated at each floor level. The 

numerical simulation did not replicate these floor slab vibrations – controlled by the local vertical 

modes – with as much precision as the horizontal response. In Figure 10-8, the numerical and 

experimental 5% damped vertical acceleration spectra at the roof level are compared for Diablo 

80% and Rinaldi 88%. In this figure, both the numerically and experimentally determined 

spectra were dominated by a single spectral peak, which occurred at the frequencies mentioned 

above, thus verifying the single mode response. However, the amplitudes of the peaks were not 

predicted accurately by the numerical simulation model. The vibration intensity at the roof level 

was significantly under-estimated and the vibration intensity at the 5th floor level was 
186

significantly over-estimated by the numerical simulation model. Although the intensity of local 

vertical modes could not be matched well by numerical simulation, this did not affect the ability 

of the model to predict horizontal floor accelerations and floor spectra in the building.

Figure 10-7 Comparison of vertical roof slab acceleration (blue), 5th floor slab 

acceleration (red), and ground acceleration (black) for (a) Vogtle 100%, (b) 

Diablo 80%, and (c) Rinaldi 88%. Peak values indicated on graph
187

Figure 10-8 Comparison of numerical simulation to experimental data for vertical slab 

acceleration; (a) roof floor and (c) 5th floor for Diablo 80%, and (b) roof floor 

and (d) 5th floor for Rinaldi 88%
188

Figure 10-9 Comparison of x and y-direction roof acceleration histories and peak 

acceleration profiles for Diablo 80% and Diablo 95% (XY)

Besides the generation of significant vertical vibration of the floor system due to vertical 

excitation, the horizontal accelerations were also amplified for the 3D records relative to the XY 

records. This phenomenon is shown in Figures 10-9 and 10-10, which compare horizontal roof 

acceleration histories and peak acceleration profiles in x and y-directions for Diablo 80% /Diablo 

95% (XY) and Rinaldi 88%/Rinaldi 88% (XY), respectively. For the Diablo record, the y-direction 

acceleration was larger for Diablo 80%, which included the vertical component of excitation, 

than for Diablo 95%, even though the scale factor applied to the motion was lower (Figure 10-9). 

The peak accelerations in the x-direction were approximately the same for Diablo 80% and 

Diablo 95% (XY). The amplification of horizontal floor acceleration under 3D excitation was even 

stronger for Rinaldi 88%. Peak horizontal accelerations were amplified by about 20% in the xdirection and about 100% in the y-direction (i.e. floor acceleration approximately doubled in the 

y-direction) (Figure 10-10). Part of this amplification was believed to be a coupled horizontalvertical mode that was driven by the vertical acceleration (Section 9.3), which will be discussed 

further in Section 10.3. For interpreting the nonstructural component response, these floor 

accelerations were larger at the roof level than at the 5th floor level.

 

10 15 20 25

-0.5

0

0.5

Acc. 6th Floor - X

Acceleration (g)

Time (sec)

10 15 20 25

-0.5

0

0.5

Acc. 6th Floor - Y

Time (sec)

0 0.5 1 1.5

Ground

Base

1st

2nd

3rd

4th

Roof

Acc. Profile - X

Peak Acceleration (g)

0 0.5 1 1.5

Ground

Base

1st

2nd

3rd

4th

Roof

Acc. Profile - Y

Peak Acceleration (g)

DIA80

DIA95_2D
189

Figure 10-10 Comparison of x and y-direction roof acceleration histories and peak 

acceleration profiles for Rinaldi 88% (XY) and Rinaldi 88%

10.3 Source of Amplification of Horizontal Floor Acceleration under 3D Input

The amplification of horizontal floor accelerations described in Section 10.2 is best understood 

by considering horizontal floor spectra, and recalling the modal analysis that was presented in 

Chapter 9. The experimentally recorded horizontal floor spectra are compared for Diablo 80% 

and Diablo 95% (XY) in Figure 10-11, and for Rinaldi 88%/Rinaldi 88% (XY) in Figure 10-12. 

Note the difference in the amplitude for the two figures. In the x-direction, the floor spectra 

peaks were not amplified for Diablo 80% relative to Diablo 95% (XY). In the y-direction, 

however, acceleration peaks were amplified significantly at about 0.18 sec in the 1st, 3rd, 4th and 

6th floors, and at about 0.1 sec in all floors, but especially 2nd, 3rd, 5th and 6th (Figure 10-11). The 

peak at 0.18 sec was associated with the 2nd structural mode (Mode 7 in Figure 9-11), and the 

peak at 0.1 sec was associated with the 3rd structural mode (Section 9.3Mode 13 in Figure 9-

13). Thus, the vertical component of acceleration amplified both the 2nd and 3rd structural 

modes.

For Rinaldi 88%, the x-direction spectral peaks at 0.18 and 0.10 sec at the associated floors 

were amplified very slightly for 3D excitation relative to XY excitation. The y-direction spectral 

peaks were amplified much more strongly at these two periods, especially the peaks at 0.18 sec 

(Figure 10-12). Thus, both 2nd and 3rd structural modes were amplified very slightly in the xdirection, while the 2nd structural mode was amplified most strongly and the 3rd structural mode 

5 10 15 20

-0.4

-0.2

0

0.2

0.4

0.6

Acc. 6th Floor - X

Acceleration (g)

Time (sec)

5 10 15 20

-1.5

-1

-0.5

0

0.5

1

Acc. 6th Floor - Y

0 0.5 1 1.5

Ground

Base

1st

2nd

3rd

4th

Roof

Acc. Profile - X

Peak Acceleration (g)

0 0.5 1 1.5

Ground

Base

1st

2nd

3rd

4th

Roof

Acc. Profile - Y

Peak Acceleration (g)

RRS88

RRS88_2D
190

also amplified significantly in the y-direction. Spectral intensities greater than 5g were observed 

in the 2nd structural mode.

The floor spectra predicted by numerical simulation for Diablo 80% and Rinaldi 88% are 

compared to the experimental data in Figures 10-13 and 10-14, respectively. For Diablo 80%, 

the numerically predicted peaks at the 2nd and 3rd structural modes were close to those 

observed from the experimental data in both x and y-directions (Figure 10-13). For Rinaldi 88%, 

the numerically predicted peaks were close to those observed from the test data for all but the 

2nd structural mode in the y-direction. (Figure 10-14).

Figure 10-11 Comparison of x and y-direction 5% damped spectral accelerations for 1st

through 6th floors for Diablo 80% and Diablo 95% (XY). These spectra were 

derived from the experimentally recorded signals

0

1

2

F1 - X F1 - Y

0

1

2

F2 - X F2 - Y

0

1

2

F3 - X F3 - Y

0

1

2

F4 - X F4 - Y

0

1

2

F5 - X F5 - Y

10-3 10-2 10-1 100

101 0

1

2

F6 - X

Period (s)

10-3 10-2 10-1 100

101

F6 - Y

Period (s)

DIA80

DIA95_2D

10-3 10-2 10-1 100

101

F6 - Y

Period (s)

10-3 10-2 10-1 100

101 0

1

2

F6 - X

Period (s)

F5 - Y

0

1

2

F5 - X

F4 - Y

0

1

2

F4 - X

Spectral Acceleration (g)

F3 - Y

0

1

2

F3 - X

F2 - Y

0

1

2

F2 - X

F1 - Y

0

1

2

F1 - X
191

Figure 10-12 Comparison of x and y-direction 5% damped spectral accelerations for 1st

through 6th floors for Rinaldi 88% (XY) and Rinaldi 88%. These spectra were 

derived from the experimentally recorded signals

The explanation for amplification of the 3rd structural mode (period = 0.1 sec) was introduced in 

Section 9.3. As described, this 3rd structural mode included substantial vertical slab vibration of 

the 4th and 6th floors, especially for the y-direction mode (mode 13 for y-direction and mode 14 

0

2

4

6

F1 - X F1 - Y

0

2

4

6

F2 - X F2 - Y

0

2

4

6

F3 - X F3 - Y

0

2

4

6

F4 - X F4 - Y

0

2

4

6

F5 - X F5 - Y

10-3 10-2 10-1 100 101 0

2

4

6

F6 - X

Period (s)

10-3 10-2 10-1 100 101

F6 - Y

Period (s)

RRS88

RRS88_2D

10-3 10-2 10-1 100 101

F6 - Y

Period (s)

10-3 10-2 10-1 100 101 0

2

4

6

F6 - X

Period (s)

F5 - Y

0

2

4

6

F5 - X

F4 - Y

0

2

4

6

F4 - X

Spectral Acceleration (g)

F3 - Y

0

2

4

6

F3 - X

F2 - Y

0

2

4

6

F2 - X

F1 - Y

0

2

4

6

F1 - X
192

for x-direction, Figure 9-14). Thus, the 3rd structural mode was driven by the vertical excitation, 

and the stronger the vertical excitation, the more this 3rd structural mode was amplified. 

Horizontal-vertical modal interaction is not a feasible explanation for the amplification of what 

appears to be the 2nd structural mode. No vertical slab vibration was observed in the 2nd

structural mode in either direction (mode 7 for y-direction and mode 8 for x-direction, Figure 9-

12). We propose two possible theories regarding the source of amplification. They are simply 

theories that cannot be confirmed by the existing evidence. The first theory is that the 2nd

structural mode was amplified by the high frequency oscillation in the base shear. Recall that a 

high frequency oscillation in the bearing shear force was clearly discernible for Rinaldi 88% 

(e.g., Figures 10-1 to 10-3), which could have been produced by the rapid change in the state of 

confinement of the lead core due to vertically induced axial load variation on the LR bearings. 

Thus, the variation of the yield force, tuned to the natural vertical frequency of the system (about 

7 Hz or 0.14 sec), introduced a small harmonic to the base shear at 0.14 sec. This harmonic 

drove the overall structure in the nearest modes, which were around 0.17 to 0.18 sec.

The following arguments support this first theory that the slight axial load induced shear force 

fluctuation contributed to the large peaks in floor spectra coincident with the 2nd structural 

modes (around 0.18 sec): (1) The numerically simulated floor spectra, which could not capture 

the aforementioned shear force fluctuation, matched the experimental floor spectra for Diablo 

80% well, since the shear force fluctuation was not evident in experimental data (Figure 10-12). 

However, the numerically simulated floor spectra fell well short of the experimental floor spectra 

at the 0.18 sec peak for Rinaldi 88% (Figure 10-12) when the shear force fluctuation was 

observed; (2) a similar amplification of the 2nd structural mode was observed in TP bearings that 

was quite evidently attributed to axial force induced fluctuations in the base shear throughout 

the experimental program (Dao 2012); (3) the amplification was strong in the y-direction (and 

essentially absent in the x-direction) for Rinaldi 88% because the input motion was much 

stronger in the y-direction than the x-direction.

The following arguments suggest that axial load induced shear force fluctuation could not have 

been singly responsible for the large peaks in the floor spectra coincident with the 2nd structural 

modes. (1) The numerical simulation model, which could not have captured the shear force 

fluctuation, predicted some amplification of the 2nd structural mode peak. (2) A base shear 

harmonic at 0.14 sec should have produced similar amplification of 2nd and 3rd structural modes, 

both about equally spaced from 0.14 sec (0.18 sec for 2nd structural mode and 0.10 sec for 3rd

structural modes), yet the 2nd structural mode seemed to be amplified much more. 
193

Figure 10-13 Comparison of numerical simulation to experimental data for Diablo 80%; x 

and y-direction 5% damped spectral accelerations for 1st through 6th 

floors. Numerical simulation uses the characterized bearing model with 

85% of the characterized value for Qd

0

1

2

F1 - X F1 - Y

0

1

2

F2 - X F2 - Y

0

1

2

F3 - X F3 - Y

0

1

2

F4 - X

Spectral Acceleration (g)

F4 - Y

0

1

2

F5 - X F5 - Y

10-3 10-2 10-1 100

101 0

1

2

F6 - X

Period (s)

10-3 10-2 10-1 100

101

F6 - Y

Period (s)

Test_X

Anal_X

10-3 10-2 10-1 100

101

F6 - Y

Period (s)

10-3 10-2 10-1 100

101 0

1

2

F6 - X

Period (s)

F5 - Y

0

1

2

F5 - X

F4 - Y

0

1

2

F4 - X

Spectral Acceleration (g)

F3 - Y

0

1

2

F3 - X

F2 - Y

0

1

2

F2 - X

F1 - Y

0

1

2

F1 - X
194

Figure 10-14 Comparison of numerical simulation to experimental data for Rinaldi 88%; 

x and y-direction 5% damped spectral accelerations for 1st through 6th

floors. Numerical simulation uses the characterized bearing model with 

85% of the characterized value for Qd. 

The foregoing considerations lead to the second theory, that amplification of a different mode 

also contributed to the amplified floor spectra peak at around 0.18 sec. Possible modes that 

0

2

4

6

F1 - X F1 - Y

0

2

4

6

F2 - X F2 - Y

0

2

4

6

F3 - X F3 - Y

0

2

4

6

F4 - X

Spectral Acceleration (g)

F4 - Y

0

2

4

6

F5 - X F5 - Y

10-3 10-2 10-1 100

101 0

2

4

6

F6 - X

Period (s)

10-3 10-2 10-1 100

101

F6 - Y

Period (s)

Test_X

Anal_X

10-3 10-2 10-1 100

101

F6 - Y

Period (s)

10-3 10-2 10-1 100

101 0

2

4

6

F6 - X

Period (s)

F5 - Y

0

2

4

6

F5 - X

F4 - Y

0

2

4

6

F4 - X

Spectral Acceleration (g)

F3 - Y

0

2

4

6

F3 - X

F2 - Y

0

2

4

6

F2 - X

F1 - Y

0

2

4

6

F1 - X
195

might have contributed are the 9th mode, which is a local vertical mode, or the 10th mode, which 

is a torsional mode with the same characteristic mode shape as the 2nd structural mode, but 

slightly more slab vibration (Figure 9-13). The 9th mode includes slight horizontal deflection and 

like the 2nd structural mode, appears to have nodes at the 2nd and 5th floor. The 9th mode 

includes a strong component of slab vibration at the roof and would thus be driven by vertical 

ground excitation.

In summary, the main source of amplification of horizontal acceleration was due to high 

frequency coupled horizontal-vertical modes of the structure that were driven by the vertical 

excitation input. A mode with frequency of 10 Hz or T = 0.1 sec was positively identified and 

other modes may have contributed. Another possible source was that axial force variation in the 

isolators introduced a high frequency component into the base shear, which drove structural 

modes around 0.18 sec.

10.4 Influence of Supplementary Roof Mass on Vertical Slab Vibration and 

Horizontal-Vertical Coupling

Being mindful of the limitations of the numerical simulation model to predict vertical slab 

vibrations, the numerically simulated vertical acceleration at the 5th floor slab and roof slab are 

compared in Figure 10-15 for the tested building model, and the augmented model with the 

supplementary weight removed from the roof. This figure predicts that relocating the roof mass 

at the base level will not much reduce the intensity of vertical slab acceleration at the roof level 

and may actually increase the slab acceleration at the 5th floor. Given that the numerical 

simulation model was not shown to predict vertical acceleration well, we do not have high 

confidence in the prediction and expect that removing the roof mass would reduce the slab 

acceleration at the roof level.

Numerical simulation predicts that removing the supplementary roof mass reduces or eliminates 

the horizontal-vertical coupling that is manifested through amplified floor accelerations in 3D 

relative to XY simulations. Figures 10-16 to 10-19 utilize the augmented model with the 

supplementary roof mass moved to the base. The numerically simulated roof acceleration 

history and peak acceleration profiles are compared for Diablo 80%/Diablo 95% (XY) (Figure

10-16) and Rinaldi 88%/Rinaldi 88% (XY) (Figure 10-17). Comparing Figure 10-6 to Figure 10-

9, almost no amplification of horizontal acceleration during 3D excitation is observed in the 

Diablo simulation when the mass was removed from the roof. The comparison is imprecise 

since the scale factors are different in the XY and 3D motions, consistent with the executed 

experiment. Comparing Figure 10-17 to Figure 10-10 (Rinaldi simulation), the peak horizontal 

floor accelerations are still amplified during 3D excitation when the mass was removed from the 

roof, but the amplification is substantially less than that observed in the experiment. These 

simulations support the theory that amplification of horizontal acceleration observed in 3D 

simulations throughout the test program are a result of horizontal-vertical coupling modes in the 

building, for which the supplementary roof mass is at least partially responsible. 
196

Figure 10-15 Numerically simulated vertical slab acceleration for Diablo 80% and Rinaldi 

88%; comparison using the testbed building model and the augmented 

model with supplementary mass removed from the roof and replaced at the 

base. 

The numerically simulated floor spectra in both x and y-directions are also compared for the 

Diablo 80%/Diablo 95% XY (Figure 10-18) and Rinaldi 88%/Rinaldi 88% XY (Figure 10-19) 

pairings. In the Diablo simulation, very slight amplification of spectral peaks at periods of 0.10 

sec and below are observed, which suggests that a small amount of horizontal-vertical coupling 

may be present in what is the equivalent of a third structural mode, which is then driven by 

vertical excitation. However, comparing Figure 10-18 to Figure 10-11, the contribution of peaks 

at periods below 0.10 sec is much less when the mass was removed from the roof, such that 

the peak horizontal floor acceleration is not affected by vertical excitation (Figure 10-18). For the 

Rinaldi simulation, almost no amplification is observed in the critical period range of 0.10 to 0.20 

sec when the mass was removed from the roof (Figure 10-19). The range of 0.10 to 0.20 sec 

comprises the range where amplification of spectral peaks in the 3D motion were significant for 

the tested building (Figure 10-11). The simulation results suggest that amplification would not be 

significant if the mass were removed from the roof, but recall that the simulation model did not 

capture all the coupling effects that were observed in the experiment. Meanwhile, the numerical 

simulation suggests that some amplification occurs at higher frequencies, or periods below 0.10 

sec. The higher frequency peaks may not be significant.

Although it cannot be confirmed without further investigation, we expect that under typical levels 

of vertical excitation (V-PGA < 0.5g) and torsional and vertical mass irregularities limited to 

accidental sources, amplification of horizontal acceleration would be negligible. Thus, the 

horizontal amplification observed in the experiments was not representative of typical practice. 

Fortunately, the numerical simulation data presented also suggests that effects of irregularities 

and amplification of horizontal acceleration can be identified during the design process.

10 12 14 16

-5

0

5

Slab Acc. 6th (g)

(a) Diablo 80%

1.81 1.78

8 9 10 11 12

-5

0

5

5.2

4.58

(b) Rinaldi 88%

10 12 14 16

-5

0

5

2.77

2.95

Time (sec)

Slab Acc. 5th (g)

8 9 10 11 12

-5

0

5

5.92

7.26

Time (sec)

Weight at Roof

Weight at Base
197

Figure 10-16 Comparison of numerically simulated x and y-direction roof acceleration 

histories and peak acceleration profiles for Diablo 80% and Diablo 95% (XY) 

using the augmented model with supplementary mass removed from the 

roof and replaced at the base 

10 15 20 25

-0.5

0

0.5

Acc. 6th Floor - X

Acceleration (g)

Time (sec)

10 15 20 25

-0.5

0

0.5

Acc. 6th Floor - Y

Time (sec)

0 0.5 1 1.5

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - X

Peak Acceleration (g)

0 0.5 1 1.5

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - Y

Peak Acceleration (g)

DIA80

DIA95_2D
198

Figure 10-17 Comparison of numerically simulated x and y-direction roof acceleration 

histories and peak acceleration profiles for Rinaldi 88% XY and Rinaldi 88% 

using the augmented model with supplementary mass removed from the 

roof and replaced at the base

5 10 15 20

-0.8

-0.6

-0.4

-0.2

0

0.2

0.4

0.6

Acc. 6th Floor - X

Acceleration (g)

Time (sec)

5 10 15 20

-0.8

-0.6

-0.4

-0.2

0

0.2

0.4

0.6

Acc. 6th Floor - Y

0 0.5 1 1.5

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - X

Peak Acceleration (g)

0 0.5 1 1.5

Ground

Base

2nd

3rd

4th

5th

Roof

Acc. Profile - Y

Peak Acceleration (g)

RRS88

RRS88_2D
199

Figure 10-18 Comparison of numerically simulated x and y-direction 5% damped 

spectral accelerations in 1st through 6th floors for Diablo 80% and Diablo 

95% (XY) using the augmented model with supplementary mass removed 

from the roof and replaced at the base

0

1

2

F1 - X F1 - Y

0

1

2

F2 - X F2 - Y

0

1

2

F3 - X F3 - Y

0

1

2

F4 - X

Spectral Acceleration (g)

F4 - Y

0

1

2

F5 - X F5 - Y

10-3 10-2 10-1 100 101 0

1

2

F6 - X

Period (s)

10-3 10-2 10-1 100 101

F6 - Y

Period (s)

DIA80

DIA95(XY)

10-3 10-2 10-1 100 101

F6 - Y

Period (s)

10-3 10-2 10-1 100 101 0

1

2

F6 - X

Period (s)

F5 - Y

0

1

2

F5 - X

F4 - Y

0

1

2

F4 - X

Spectral Acceleration (g)

F3 - Y

0

1

2

F3 - X

F2 - Y

0

1

2

F2 - X

F1 - Y

0

1

2

F1 - X
200

Figure 10-19 Comparison of numerically simulated x and y-direction 5% damped spectral 

accelerations in 1st through 6th floors for Rinaldi 88% XY and 3D using the 

augmented model with supplementary mass removed from the roof and 

replaced at the base

0

1

2

F1 - X F1 - Y

0

1

2

F2 - X F2 - Y

0

1

2

F3 - X F3 - Y

0

1

2

F4 - X

Spectral Acceleration (g)

F4 - Y

0

1

2

F5 - X F5 - Y

10-3 10-2 10-1 100 101 0

1

2

F6 - X

Period (s)

10-3 10-2 10-1 100 101

F6 - Y

Period (s)

RRS88

RRS88(XY)

10-3 10-2 10-1 100 101

F6 - Y

Period (s)

10-3 10-2 10-1 100 101 0

1

2

F6 - X

Period (s)

F5 - Y

0

1

2

F5 - X

F4 - Y

0

1

2

F4 - X

Spectral Acceleration (g)

F3 - Y

0

1

2

F3 - X

F2 - Y

0

1

2

F2 - X

F1 - Y

0

1

2

F1 - X
201

10.5 Implications for Design

This investigation of the response of the testbed structure for 3D excitations has shown that the 

influence of vertical excitations on the response of the seismic response of the system can be 

significant. Substantial vertical slab vibrations were generated, and horizontal floor accelerations 

were amplified by the input vertical motions. Fortunately, these effects were replicated by a 

properly crafted numerical simulation model of the structure and the isolators. The testbed 

structure was not best configured to minimize the effects of vertical excitation. Thus, the results 

have reinforced that good design practices should always be followed. Site specific horizontal 

and vertical acceleration histories should be used in the analysis and design of the structure, as 

is already done routinely for nuclear structures but is not required for conventional (non-nuclear) 

structures. The design of the base isolation system should consider and accommodate any 

sources of coupling in the structure, including horizontal-torsional and horizontal-vertical 

coupling effects.

203

11. STRUCTURAL AND ISOLATOR RESPONSE FOR THREE BUILDING 

CONFIGURATIONS

In this chapter, the general responses of the testbed building and isolation system are 

compared for the hybrid LR isolation system configuration, the TP isolation configuration, and 

the fixed-base configuration. When comparing the two isolation configurations, we emphasize 

that differences in isolator displacement demands and base shear are primarily determined by 

the design properties of the isolators and not related to the type of device. The normalized 

backbone curves (normalized base shear vs. isolator displacement) are compared for the hybrid 

LR isolation system and the TP isolation system in Figure 11-1. This figure is based on the 

design parameters of the hybrid LR isolation system (Table 3-3) rather than the characterized 

parameters. The normalized strength Qd/W and post-yield period Td represent the entire 

isolation system (i.e. strength and stiffness summed over all bearings); they are calculated using 

the measured weight of the testbed structure (Section 4.3) and thus differ from the design 

values given in Chapter 3. The backbone curve of the TP isolation system also reflects design 

parameters, wherein the multi-linear curve reflects the friction coefficients and pendulum lengths 

of the bearing’s independent pendulum mechanisms. For more information about the design, 

refer to Dao (2012).

Figure 11-1 Normalized backbone curves (base shear vs. displacement) for the Long 

Period TP isolation system and the Moderate Period LR isolation system, 

based on design properties of the bearings and measured mass

With a post yield period Td = 2.8 sec, the as-designed hybrid LR isolation system provides a 

moderate period shift. In contrast, the TP isolation system when fully engaged (Td = 5.6 sec) 

provides a very long period shift. For the same seismic hazard, the lengthening of the isolation 

period is expected induce larger displacement demands, and as such the TP isolation system 

was designed with a large displacement capacity greater than 1 meter. While an LR isolation 

system could be designed to provide a backbone curve comparable to the tested TP isolation 

system, we remind the reader that providing this combination of flexibility and displacement 

capacity in a system designed with elastomeric or LR bearings might be challenging. On the 

other hand, significant penalties for large displacements would be incurred in the design of a 

safety related nuclear facility, and thus the isolation period of a nuclear structure is likely to be 

less than 3 sec. To emphasize that the response depends on the design parameters, we refer to 
204

the hybrid LR isolation system as the “Moderate Period” system and the TP isolation system as 

“Long Period” system. However, some response trends are device specific and will be identified 

as such.

Several records were applied commonly to all three configurations to allow their responses to be 

compared objectively. Although the complete simulation schedule was summarized in Tables 4-

3 to 4-5, the comparative simulations were not identified. Thus, the applied records, scale 

factors, target PGA and realized table PGA for all comparative simulations are summarized in 

Table 11-1. The Westmorland record (WSM80) from the 1987 Superstition Hills Earthquake was 

intended to represent a service level earthquake for the TP isolation configuration, Dao 2012) 

and was scaled to 80%. WSM80 is the only record that was applied to the three configurations 

at the same scale factor. The Iwanuma record (Iwanuma 100% (XY)) is a long period, long 

duration (approximately 200 sec) subduction record recorded in the 2011 Tohoku Earthquake, 

and was selected by our Japanese collaborators as it is typical of the seismicity in Japan. The 

record was scaled to 70% for the fixed-base building, which, as explained earlier, was selected 

to avoid any possibility of yielding in the moment frame structure. The vertical component of 

Iwanuma was not included since this would have precluded simulating the entire duration of the 

record due to limitations of the earthquake simulator. The Rinaldi record (Rinaldi 88%) is a wellknown near-fault record from the 1994 Northridge Earthquake, which was scaled to 88% to 

represent a design level earthquake for the TP isolation configuration (Dao 2012). As explained 

earlier, this record was executed without the vertical component – denoted XY – and with the 

vertical input to assess the influence of vertical excitation in all three systems. The influence of 

vertical excitation for the hybrid LR configuration was examined thoroughly in Chapter 10; so 

that this chapter highlights the comparison between the systems. Note that the horizontal 

components of the Rinaldi record were scaled down to 35% for the fixed-base configuration, for 

the same reason described above, but the target vertical component was the same for all three 

system configurations. Recall that the realized peak accelerations of the earthquake simulator 

were generally larger than the target accelerations (Section 4.5), which is also observed in 

Table 11-1.

Table 11-1 Ground Motions, Scale Factors and Peak Accelerations Common to 3 

Configurations. 

Ground motion System

Target 

Scale 

Factor

Target 

Peak 

Axy (gg)

Realized 

Peak

Axy (g) 

Target 

Peak 

Az (gg)

Realized 

Peak

Az (gg)

1987 Superstition Hills

Westmorland

Hybrid 80% 0.17 0.19 0.20 0.14

TP 80% 0.17 0.18 0.20 0.15

Fixed 80% 0.17 0.24 0.20 0.14

2011 Tohoku

Iwanuma (XY)

Hybrid 100% 0.43 0.59 0 0.02

TP 100% 0.43 0.60 0 0.03

Fixed 70% 0.30 0.39 0 0.01

1994 Northridge

Rinaldi Rec. Sta. (XY)

Hybrid 88% 0.72 1.18 0 0.05

TP 88% 0.72 1.24 0 0.10

Fixed 35% 0.28 0.41 0 0.01

1994 Northridge

Rinaldi Rec. Sta.

Hybrid 88% 0.72 1.19 0.73 1.25

TP 88% 0.72 1.26 0.73 1.26

Fixed 35% xy 

88% z 0.28 0.42 0.73 1.10
205

11.1 Structural Response

Acceleration profiles (peak floor acceleration versus floor level) are compared in Figure 11-2 for 

each building configuration and each record. The accelerations plotted represent the peak 

instantaneous vector sum of floor acceleration at the center of the building (average over the 

NW and SE sensors). For a clearer impression of the level of attenuation or amplification of 

each building systems, the accelerations are normalized by PGA (different for each 

configuration) on the right graph. For the first three records (Figure 11-2(a) to (c)), the 

acceleration in the isolated building configurations was attenuated well below PGA – as 

expected – and the acceleration was essentially constant over the height. On the other hand, 

the acceleration increased with height in the fixed-base configuration, where at the roof level, 

the acceleration was amplified by a factor of about 2.5 to 3 with respect to PGA. 

For the low level WSM80 excitation, the accelerations were better attenuated for the Moderate 

Period isolation system than for the Long Period isolation system, while for the larger intensity 

Iwanuma and Rinaldi XY excitations, the opposite was true. An explanation of this behavior is 

given in Section 11.2. For both systems, the attenuation of acceleration increased as the 

intensity of the record increased. Thus, the peak floor accelerations were about 60-70% of PGA 

in WSM80, 40-50% of PGA in Iwanuma, and 20-30% of PGA in Rinaldi XY. Thus, for Rinaldi 

XY, the acceleration at the roof level was reduced by about a factor of 10, which is compelling 

evidence of the effectiveness of base isolation. 

The influence of a strong vertical excitation for all configurations can be ascertained by 

comparing the responses of each system observed during Rinaldi 88% to those observed 

during Rinaldi 88% (XY) (Figures 11-2(c) and (d)). The influence of vertical excitation was 

influenced by the type of isolation device and thus will be described with reference to the 

devices. To summarize observations for the hybrid LR system that have already been 

described, significant amplification of horizontal acceleration was observed in Rinaldi 88% (3D 

excitation) compared to Rinaldi 88% (XY) (Section 5.6), and the potential sources of 

amplification were identified (Section 10.3). The sources of amplification were: 1) high frequency 

coupled horizontal-vertical modes of the structure that were driven by the vertical excitation 

input (a mode with frequency of 10 Hz or T = 0.1 sec was positively identified), and 2) axial 

force variation in the isolators introduced a high frequency component into the base shear, 

which drove structural modes around 0.18 sec. Based on Figure 11-2(c) and (d), the 

acceleration was also amplified in the fixed-base configuration for the Rinaldi 88% excitation; 

the roof acceleration increased from about 1g for XY input to 1.25g for 3D input. Thus, 

amplification of acceleration was noteworthy in the fixed-base configuration, but not as strong as 

for the hybrid LR isolation configuration since only the first source of amplification (modal 

coupling) would be present in the fixed-base building. As discussed earlier, the amplification of 

horizontal acceleration in the 3D excitation was largely a result of the supplementary roof mass, 

and was reduced significantly when the mass was removed from the roof in the numerical 

simulation model (Section 10.4). 
206

Figure 11-2 Peak floor acceleration (vector sum) vs. floor level of the three building 

configurations for (a) WSM 80%, (b) IWA 100%/70%, (c) Rinaldi 88%/35% 

(XY) and (d) Rinaldi 88%/35% records. Accelerations are absolute on the 

left graph and normalized by PGA on the right. 

0 0.2 0.4 0.6

Table

Base

2

3

4

5

Roof

Floor Level

(a) WSM 80%

0 1 2 3

0 0.5 1 1.5

Table

Base

2

3

4

5

Roof

Floor Level

(b) Iwanuma 100%,70%

0 1 2 3

0 0.5 1 1.5

Table

Base

2

3

4

5

Roof

Floor Level

(c) Rinaldi 88%,35% XY

0 1 2 3

0 0.5 1

Table

Base

2

3

4

5

Roof

Floor Level

Peak Accel. (g)

(d) Rinaldi 88%,35% 3D

0 1 2 3

Peak Accel./PGA

TPB Hybrid Fixed
207

As observed in Figure 11-2, the amplification of horizontal accelerations was greater for the 

Long Period TP isolation system than for the Moderate Period hybrid LR isolation system. The 

peak floor accelerations, which were lowest for Rinaldi XY in the TP isolation configuration as a 

result of its long period shift (Figure 11-2(c)), were generally larger the TP isolation configuration 

than the hybrid LR isolation configuration for Rinaldi 3D (Figure 11-2(d)). The amplification was 

larger for TP bearings (a device specific effect) because an additional source of coupling is 

present in the bearings. In TP bearings or any friction devices, the horizontal force (base shear) 

is directly proportional to the vertical force (or axial load) carried by the bearing. Thus, the total 

axial force variation on the isolators due to vertical excitation is echoed as a high frequency 

oscillation in the total base shear. To demonstrate, the history of total base shear coefficient of 

the three building configurations for Rinaldi 88% (XY) and Rinaldi 88% motions is illustrated in 

Figure 11-3. For Rinaldi 88% (XY), the base shear of the isolated configurations oscillated at the 

natural frequency of the system, which was similar for the two configurations. The peak base 

shear of the hybrid LR isolated building was essentially unaffected by the vertical excitation 

(Figure 11-3(b)). On the other hand, the peak base shear coefficient of the TP isolated building 

increased from V/W = 0.13 for XY input to V/W = 0.24 for 3D input. This increase was caused 

by the high frequency oscillation observed in the base shear (Figure 11-3(b)). This oscillation 

has been confirmed to coincide with the fundamental vertical frequency of the structure, which 

was controlled by slab vibration at the roof level (similar to mode 9 in Figure 9-13). 

Introduction of an additional frequency component into the base shear was offered as an 

explanation for the amplification of floor spectra peaks corresponding to the second structural 

mode in the hybrid LR isolation configuration under Rinaldi 88% (Section 10.4). However, the 

additional frequency component was quite small for the LR system, and is not even visually 

detectable in Figure 11-3. For many motions, the additional frequency component introduced to 

the base shear was strong for the TP isolation system, and the typical pattern was that the 

second structural mode of the TP isolated building, being close to the driving vertical frequency, 

was driven strongly by the base shear. However, because the vertical excitation in Rinaldi 88% 

(> 1g) was strong enough to cause near complete uplift of the building from the isolators, the 

response of the system was complex. The 5% damped floor spectra for the TP isolation 

configuration observed in Rinaldi 88% (XY) and Rinaldi 88% (Figure 11-4) demonstrate the 

amplification of higher modes of the system over a range of frequencies. The reader is referred 

to Dao (2012) for a full documentation of the influence of vertical excitation in TP bearings. The 

results shown in Figures 11-2(d) and 11-4 are noted to be quite sensitive to the filtering 

parameters; a Butterworth filter with a cutoff frequency of 50 Hz was utilized.

Even under the influence of vertical excitation, the base shear of both isolated configurations 

was still significantly less than the base shear of the fixed-base configuration, which was 

subjected to only about 40% of the horizontal loading of the isolated configurations (ground 

motion scale factor of 35% compared to 88%). The base shear of the fixed-base structure was 

also sensitive to a high frequency component under Rinaldi 88% (Figure 11-3(b)) that was not 

present under Rinaldi 88% (XY) (Figure 11-3(a)). The comparison of the fixed-base and isolated 

building base shear was affected by the method used to derive the base shear in the respective 

configurations. In the fixed-base building, the base shear was derived from inertial forces (the 

sum of mass multiplied by floor accelerations) while the base shear was derived from the load 

cell measurements in the isolation configurations. When applied to the isolated buildings, the 

inertial force method was confirmed to agree closely with the load cell data for XY excitations, 

but deviate from the load cell measurements as the vertical intensity increased.
208

Figure 11-3 History of total base shear in x and y-directions compared for the 3 system 

configurations in (a) Rinaldi 88% (XY) and (b) Rinaldi 88%. The bottom 

graphs provide a close-up of the isolated configurations; peak values are 

indicated

5 10 15 20

-0.6

-0.3

0

0.3

0.6

-0.3

0

0.3

-0.30

-0.67

Base Shear Coefficient

(a) Rinaldi 88% (XY)

5 10 15 20

-0.2

0

0.2

-0.2

0

0.2 0.13

0.21

0.13

0.24

Time (sec)

5 10 15 20

-0.32

(b) Rinaldi 88%

-0.72

Fixed Base TPB System Hybrid System

5 10 15 20

-0.22

0.21

0.24 0.23

Time (sec)

x-dir

y-dir

x-dir

y-dir y-dir

y-dir

x-dir x-dir
209

Figure 11-4 Comparison of x and y-direction 5% damped spectral accelerations in the 

TP isolation configuration for 1st through 6th floors for Rinaldi 88% XY and 

Rinaldi 88%

0

2

4

F1 - X F1 - Y

0

2

4

F2 - X F2 - Y

0

2

4

F3 - X F3 - Y

0

2

4

F4 - X

Spectral Acceleration (g)

F4 - Y

0

2

4

F5 - X F5 - Y

10-3 10-2 10-1 100

101 0

2

4

F6 - X

Period (s)

10-3 10-2 10-1 100 101

F6 - Y

Period (s)

10-3 10-2 10-1 100 101

F6 - Y

Period (s)

10-3 10-2 10-1 100 101 0

2

4

F6 - X

Period (s)

F5 - Y

0

2

4

F5 - X

F4 - Y

0

2

4

F4 - X

Spectral Acceleration (g)

F3 - Y

0

2

4

F3 - X

F2 - Y

0

2

4

F2 - X

F1 - Y

0

2

4

F1 - X

RRS88-TPB RRS88(XY)-TPB
210

Besides the amplification of horizontal acceleration, the recorded vertical accelerations in the 

structure were significant in all three system configurations under Rinaldi 88% 

(RRS35(XY)88(Z) for the fixed-base structure). Figure 11-5 compares the input ground 

acceleration, column accelerations, and slab accelerations in the vertical direction at the 5th and 

roof level floors in all three system configurations. The peak ground acceleration in all three 

systems was similar with peak amplitude of about 1.2g. 

In the TP configuration, a peak vertical acceleration of about 4g was observed in the columns 

(Fig 11-5(a),(d)), whereas only mild amplification of the vertical ground acceleration was 

observed in the hybrid LR isolation and fixed-base configurations (Figure 11-5(b-c),(e-f)). The 

source of the amplification in the TP isolated building was a vertical motion induced uplift that 

caused the entire building to lift off the bearings multiple times; the large accelerations were 

generated by the impact force that resulted when the building re-established contact with the 

simulator platform. Evidence of this uplift was observed in the force data for the bearings 

(simultaneous zero forces recorded in all but the SW bearing), but not shown here for brevity. 

However, the slab acceleration was dominated by vibration at its natural frequency (about 7 Hz 

at the roof level), and any high frequency input at the column level was not transmitted to the 

slabs. Thus, the amplitude of slab vibration was similar in all three configurations, and did not 

appear to be amplified in the isolation configurations over the fixed-base configuration except 

perhaps for some isolated spikes.

Figure 11-5 Vertical acceleration history of floor slab (recorded at middle of the NE 

quadrant), column average, and ground acceleration average (recorded at

simulator platform) for Rinaldi 88% for (a,d) TP isolation configuration, (b,e) 

hybrid LR isolation configuration, and (c,f) fixed-base configuration; (a-c) = 

5th floor slab and (d-f) = roof slab

Drift profiles (peak story drift vs story level) are compared in Figure 11-6 for each building 

configuration and each record. The peak drift represents the maximum relative story 

displacement in the x or y-directions at the center of the building determined by linear 

interpolation of the NW and SE displacement transducers, which directly measured relative 

story displacement. In the second story and above, a story drift of 3 cm corresponds to a drift 

angle of 0.01 rad. Because the 1st story was taller than 2nd to 5th stories (Figure 2-2), the peak 

drift angle was generally largest in the 2nd story and decreased substantially in the upper floors 

of the building. Substantial reductions in drift were observed in the isolated configurations 

0 2 4 6 8

-6

-3

0

3

6

9

Accel. (g)

(a)

Floor Slab Column Ground

0 2 4 6 8

-6

-3

0

3

6

9

(b)

0 2 4 6 8

-6

-3

0

3

6

9

(c)

0 2 4 6 8

-6

-3

0

3

6

9

Accel. (g)

(d)

Time (s)

0 2 4 6 8

-6

-3

0

3

6

9

(e)

Time (s)

0 2 4 6 8

-6

-3

0

3

6

9

(f)

Time (s)
211

relative to the fixed-base configuration, recalling the reduced scale factors to the input motions 

for IWA 100% (70% for fixed-base, Figure 11-6(b)) and Rinaldi 88% XY (35% for fixed-base, 

Figure 11-6(c)). The drift demands in the hybrid LR system (Moderate Period) and TP system 

(Long Period) were about the same in Westmorland, but became increasingly lower in the 

hybrid LR system relative to the TP system as the intensity of the motion increased (e.g. Rinaldi 

88% XY in Figure 11-6(c)), which was solely a function of the design properties of the system 

and not the specific type of isolation device. Unlike the floor acceleration, the peak drift was not 

too sensitive to vertical excitation (compare Figure 11-6(d) to Figure 11-6(c)), as the drift 

demands were determined by the fundamental or low frequency modes of the building. 

Figure 11-6 Peak story drift (max in x or y-direction) vs. story level of the three building 

configurations for (a) WSM 80%, (b) IWA 100%/70%, (c) Rinaldi 88%/35% 

(XY) and (d) Rinaldi 88%/35% records

11.2 Isolation System Response

As described above, the Long Period (TP) isolation system was designed with a displacement 

capacity of more than 1 meter. Several ground motions were selected that were intended to 

develop displacement demands close to the displacement limit of the isolation system. 

However, the peak displacement demand observed in the system was about 70 cm (27.6 in) 

under Tabas 100% (XY). The lower than projected displacement demands resulted from 

increased friction in the system; the measured friction coefficient for the outer pendulum 

mechanism (~0.10) exceeded the design friction coefficient (~0.08) for the outer pendulum 

mechanism. Besides the peak displacement demand, the suite of 19 ground motions (excluding 

sine waves) applied to the Long Period isolation system induced an average displacement 

demand of 42 cm (16.5 in), while the average displacement demand for the Moderate Period 

isolation system was 34 cm (13.4 in). Thus, the average displacement demands were not 

appreciably larger for the Long Period isolation system. This is not a fair comparison, since the 

suite of motions applied to the two systems contained little overlap. However, we observed that 
212

for some of the near-fault motions, such as Takatori 100%, a peak in spectral displacement was 

observed in the 2-3 second range, such that further lengthening of the isolation system reduced 

the expected displacement demand. Takatori 100% induced a displacement demand of 56 cm 

(22 in) in the Long Period isolation system, while Takatori was not simulated for the Moderate 

Period isolation system, as it was predicted to exceed the safety limit. Thus, if justified by the 

seismicity, the period lengthening strategy used in the design of the Long Period isolation 

system might be used to increase the factor of safety in the design (provide extra displacement 

capacity) without much increase in the expected displacement demands.

The detailed responses of the Moderate Period (hybrid LR) isolation system and Long Period 

(TP) isolation system are compared for two excitations, WSM 80% and Rinaldi 88% (XY). The 

discussion is supported by the following figures, all of which directly compare the observed 

responses in the two systems: isolator displacement traces and displacement histories (Figure

11-7 for WSM 80% and Figure 11-8 for Rinaldi 88% (XY)), isolator force vs. displacement or 

hysteresis loops (Figure 11-9), and base rotation angle (Figure 11-10).

For WSM 80%, the peak vector sum isolator displacement, indicated on the displacement trace 

(Figure 11-7), was 8.8 cm for the Moderate Period system and 15.6 cm for the Long Period 

system, both occurring in the East bearing. The difference, however, was due primarily to a 

large residual displacement in the Long Period isolation system (-6 cm in x and 1 cm in y) that 

was present at the beginning of the record. The displacement traces and histories of the two 

isolation systems were similar. Visually, the two isolation systems appeared to be responding at 

the same or similar frequencies, as the larger displacement cycles were in phase throughout the 

record. The peak in both systems was determined by a circular orbit in the displacement trace. 

The Long Period (TP) system moved nearly exclusively on the inner sliders during WSM 80%, 

as indicated by the hysteresis loops for the East bearing (Figure 11-9(a)). As such, the isolation 

system did not re-center (TP bearings must be driven to the outer part of the hysteresis loop to 

re-center). For motion on the inner sliders, the TP system was stiffer than the Moderate Period 

LR isolation system (with a tangent period of 1.8 sec compared to 2.6 sec), and had a higher 

yield force coefficient, observed to be V/W ~ 0.10 (Dao 2012, Dao et al. 2012a). On the other 

hand, the hybrid LR system yielded at a base shear coefficient V/W ~ 0.05, and thus engaged 

the softer second slope of the LR bearings. As a result, the Long Period TP system experienced 

larger base shear, and larger floor accelerations.

As the intensity of the input excitation increased, the outer pendulum mechanisms of the Long 

Period TP isolation system, with a tangent period of 5.6 sec, were engaged, and subsequently 

the base shear demand increased very slowly. This situation is well demonstrated by the 

response during Rinaldi 88% (XY). The peak vector sum displacement in the two isolation 

systems was comparable (38.8 cm or 15.3 in for the Moderate Period system and 36.6 cm or 

14.4 in for the Long Period system – both in the South bearing, Figure 11-8). This displacement 

occurred at about 9 sec, which corresponded to the timing of the near-fault fling pulse (Figure 4-

18). As expected, the base shear demand in the Moderate Period system at this displacement 

was larger than in the Long Period system (compare the hysteresis loops in Figure 11-9(b)), 

which affected the floor accelerations and story drifts, as discussed earlier. Aside from the 

instant of the large pulse, the displacement demand of the Long Period TP system was less 

than the Moderate Period LR system throughout the majority of the record (Figure 11-8), which 

suggests that the motion of the Long Period TP system was again dominated by sliding on the 

inner sliders.

Base rotation demands were affected by the isolation devices; however, as mentioned several 

times, the torsional properties of the tested hybrid LR isolation system were not representative 

of common practice, where it would be possible to configure the isolation system to minimize 

the eccentricities that induce torsion. The peak base rotation (Figure 11-10(a)) was larger in the 
213

Long Period TP system for WSM 80%, which was primarily a result of the initial residual 

displacement. On the other hand, the oscillation of rotation angle was greater in the Moderate 

Period LR system, though not particularly large for this motion. The base rotation of the hybrid 

LR system during Rinaldi 88% (XY) was substantially larger than in the TP system, which was 

typical of many of the motions. As discussed earlier, the peak rotation angle of the hybrid LR 

system in any motion was 0.019 rad, which was observed in Vogtle 175%, while the peak 

rotation angle of the TP system in any motion was 0.005 rad.

Since the shear force in a TP bearing or any friction bearing is proportional to the axial force, the 

center of resistance naturally aligns with the center of mass and rotational demands are 

expected to be negligible. The observed rotation in the TP system was small, but not negligible. 

The distribution of gravity loads on the bearings did not follow the natural weight distribution in 

the building, but rather was constrained by the rigidity of the base diaphragm. The base 

diaphragm was believed to be warped by constant exposure to weather for more than 2 years 

before the tests. The misfit induced by the warping was accommodated by pre-stress, which 

ultimately caused the weight on the isolators to be redistributed. The expected and actual 

weights carried by the TP isolators are compared in Figure 11-11. The balance of loads 

depicted was achieved only after several iterations with shimming plates. Thus, the forces 

generated in the isolators did not correspond to the actual mass distribution in the building, 

causing a moderate torsional response to be generated. The rotational input observed in the 

earthquake simulator was negligible.
214

Figure 11-7 Comparison of Moderate Period LR and Long Period TP systems for WSM 

80%; displacement trace and displacement histories (x and y) in each 

bearing, labeled by position (E, S, N, W)

-15

-10

-5

0

5

10

15

Displacement Trace

8.8

E 15.6

-10

-5

0

5

10

-5

0

5 X

Y

E

Displacement History

-15

-10

-5

0

5

10

15

7.5

S 13.4

-10

-5

0

5

10

-5

0

5 X

Y

S

-15

-10

-5

0

5

10

15

7.8

N 13

- X Displacement (cm)

-10

-5

0

5

10

-5

0

5 X

Y

N

Displacement (cm)

-15-10 -5 0 5 10 15

-15

-10

-5

0

5

10

15

7.3

11.2 W

Y Displacement (cm)

5 10 15 20 25 30 35

-10

-5

0

5

10

-5

0

5 X

Y

W

Time (sec)

Hybrid TPB
215

Figure 11-8 Comparison of Moderate Period LR and Long Period TP systems for Rinaldi 

88% (XY); displacement trace and displacement histories (x and y) in each 

bearing, labeled by position (E, S, N, W)

-40

-20

0

20

40

Displacement Trace

38.3

36.4

E

-30

0

30

-30

0

30

X

Y

E

Displacement History

-40

-20

0

20

40

38.8

36.6

S

-30

0

30

-30

0

30

X

Y

S

-40

-20

0

20

40

35.7

35.8

N

- X Displacement (cm)

-30

0

30

-30

0

30

X

Y

N

Displacement (cm)

-40 -20 0 20 40

-40

-20

0

20

40

36.5

36.1

W

Y Displacement (cm)

5 10 15 20

-30

0

30

-30

0

30

X

Y

W

Time (sec)

Hybrid TPB
216

Figure 11-9 Comparison of the base shear coefficient for Moderate Period LR and Long 

Period TP systems in the x- and y-directions for the East bearing; WSM 80% 

and Rinaldi 88% (XY)

Figure 11-10 Rotation angle at the base for Moderate Period LR and Long Period TP

isolated configurations for (a) WSM 80% and (b) Rinaldi 88% (XY)

-20 -10 0 10 20

-0.2

-0.1

0

0.1

0.2 East Bearing - X

(a) Force vs Disp. - WSM80

Displacement (cm)

Base Shear Coefficient

-20 -10 0 10 20

-0.2

-0.1

0

0.1

0.2 East Bearing - Y

Displacement (cm)

-40 -20 0 20 40

-0.4

-0.2

0

0.2

0.4 East Bearing - X

(b) Force vs Disp. - RRS2D

Base Shear Coefficient

Displacement (cm)

-40 -20 0 20 40

-0.4

-0.2

0

0.2

0.4 East Bearing - Y

Displacement (cm)

Hybrid TPB

0 10 20 30 40

-0.01

0

0.01

0.0042

0.0029

Time (s)

Rotation angle (rad)

(a) Westmorland 80%

Hybrid System

TPB System

0 5 10 15 20 25

-0.01

0

0.01

0.0032

0.016

(b) Rinaldi 88% XY

Time (s)
217

Figure 11-11 Qualitative illustration of expected weight distribution (calculated by 

natural distribution of the weight) and observed weight distribution 

(measured in load cells) on TP isolation system

The accumulation of residual or permanent displacements is distinctly a function of isolation 

period shift alone and completely independent of the isolation device. The bearing displacement 

histories shown in Figures 11-7 and 11-8 suggest that the Long Period isolation system was 

subjected to larger residual displacements than the Moderate Period isolation system. For both 

systems, the largest measured residual displacement over all the bearings after every trial is 

plotted in Figure 11-12, which confirms the observation. The largest residual displacement 

observed in the Moderate Period isolation system throughout the test program was about 2 cm 

(0.8 in), but as discussed in Section 5.5, was affected by slippage of the base plates so that the 

actual residual displacement was only about 0.5 cm (0.2 in). This residual displacement is small 

relative to the peak displacement of the system (55 cm (21.7 in) in Diablo 95%), and barely 

discernible on a cumulative displacement history of a bearing over the course of a single test 

day (Figure 11-13). On the other hand, the largest residual displacement observed in the Long 

Period isolation system throughout the test program was close to 11 cm (4.3 in), which is 

significant relative to the peak displacement of 70 cm (27.6 in). These residual displacements 

are a consequence of the flexibility of the system; the system was designed with a tangent 

period of 5.6 sec for motion on the outer pendulum mechanism. 

0

500

1000

1500

0

200

400

600

800

1000

Axial Force (kN)

Axial Force (kN)

Expected Weight 

Distribution

Observed Weight Distribution

Load on center bearing 

large

Load on center bearing 

small
218

Figure 11-14 illustrates the cumulative displacement history of a bearing in the Long Period 

isolation system over the course of a test day. Analysis of this figure suggests that the residual 

displacements did not accumulate. Near-fault motions (LGPC, ChiChi and Takatori) caused 

large residual displacements. Long duration motions or motions with typical frequency content 

(Iwanuma) helped re-center the bearings. When ChiChi 70% (denoted TCU in Table 4-3) was 

followed by ChiChi 80%, the residual displacement was larger in ChiChi, 80% but only because 

the peak displacement was larger in ChiChi 80%. The numerical simulation of the Long Period 

system (Dao 2012) suggested that these residual displacements were predictable. Furthermore, 

to reproduce the experimental histories through numerical simulation, it was not necessary and 

was in fact disadvantageous to zero correct the data. The experimental displacement histories 

(starting from the residual displacement) generally became aligned with the numerically 

simulated displacement histories (starting from zero) within the first few cycles of motion.

Although better control of the base shear demand is attractive, the accumulation of significant 

residual displacement is a noteworthy disadvantage of the Long Period design strategy. Recentering the isolation system after a large earthquake could be problematic. An aftershock 

could help to re-center the system, provided it was strong enough to engage the post-yield 

period of the system (outer most sliding mechanism for TP bearings). The capacity of the 

structure to tolerate residual displacements must be considered when employing a Long Period 

design strategy utilizing large displacement capacity. In a nuclear facility, the structure could be 

jacked off the hard stops to eliminate the permanent displacement. 

Figure 11-12 Maximum residual displacement in the Long Period and Moderate Period 

isolation systems after each simulation

Long Period System Moderate Period System

Peak Residual Disp. = 10.8 cm Peak Residual Disp. = 2.01 cm

Average Residual Disp. = 5.4 cm Average Residual Disp. = 1.2 cm
219

Figure 11-13 Displacement of the SW bearing in the Moderate Period isolation system 

over the course of a single test day, illustrating the residual displacement 

after each record

Figure 11-14 Displacement of the SE bearing in the Long Period isolation system over the 

course of a single test day, illustrating the residual displacement after each 

record

221

12. CONCLUSIONS AND CONSIDERATIONS FOR NUCLEAR POWER 

PLANTS

As documented in this report, a hybrid elastomeric isolation system using lead-rubber (LR) 

bearings and cross-linear (CL) bearings was designed for a 5-story moment frame building and 

tested under a variety of earthquake excitations at E-Defense. The isolation system was 

designed to sustain displacement demands representative of extended or beyond design basis 

shaking at a potential nuclear site in central and eastern U.S., and such motions were prioritized 

in developing the test program. The experimentally observed response of the LR bearings was 

calibrated to a bidirectionally coupled, bilinear hysteretic model with uncoupled response in the 

horizontal and vertical directions. A realistic numerical simulation model of the 5-story building 

with isolators was built, tested and calibrated. This report has documented the overall test 

results, unusual response characteristics of the hybrid isolation system, the comparison of the 

numerical simulation to experimental data, the influence of vertical excitation, and the 

comparison of the hybrid LR isolation system to two other building configurations that were also 

tested. Key observations and considerations for nuclear power plants are discussed as follows.

12.1 Characteristics of Lead-Rubber Bearing Response

The following behaviors, many of which have been observed before, were observed in the 

response of LR bearings during this test program. 

1. Pinching near the center of the measured bearing hysteresis loop, attributed to the small 

size of the lead plug;

2. Loss of characteristic strength over the duration of an excitation, associated with heating of 

the lead plug;

3. Slight fluctuation of shear force during high frequency axial force variation; thought to be 

insignificant;

4. Small (negligible) permanent displacements at the end of the records;

5. Significant base rotation demands due to the inability to configure the system appropriately 

for torsion.; 

6. No loss of shear resistance at large displacements due to the stabilizing influence of the CL 

bearings;

7. Transfer of axial forces from LR bearings to CL bearings at large displacements, causing the 

LR bearings to sustain tension;

Items 1-5 are not believed to be influenced by the presence of CL bearings. With regard to item 

6, the stabilizing influence of the CL bearings prevented the loss of shear resistance of the LR 

bearings at displacement demands beyond their theoretically computed stability limits in this 

experiment. Normally, a system composed only of LR bearings can be designed to stay well 

within the stability limits, and under this scenario similar behavior would be expected. Item 7 is a 

behavior unique to the hybrid LR system.

12.2 Predictability of the System Response

According to the recently developed NUREG (Kammerer et al. 2012), Predictability of response 

is perhaps the most has been identified as an important requirement for the application of 
222

seismic isolation to nuclear power plants. Our numerical simulation and its comparison to 

experimental data cannot be directly applied to nuclear facilities due to difficulties in modeling 

the response of the tested LR bearings, and differences between the tested superstructure and 

a nuclear facility. Nonetheless, some of the observations and insights learned during this 

simulation exercise are applicable to isolated structures in general.

The following should be noted when evaluating the outcome of our numerical simulation to 

match the experimental data. A single amplitude-independent model with parameters based on 

physical theory is desirable to represent the response of the isolation system. Such an approach 

might be possible in general, but was not possible in this study due to the pinching of the 

hysteresis loops near zero displacement, which was a result of the small size of the lead plug. 

The pinching effect would greatly diminish or disappear for bearings with lead plugs designed to 

a minimum recommended diameter equal to at least 1/6 the bonded diameter of the bearing.

Second, the bearings were modeled using a rate independent bidirectionally coupled plasticity 

model that is commonly available in numerical simulation software. Improved models are 

available that can capture variation of the yield strength due to heating of the lead plug or 

pressure variation due to cyclic axial loading; both of these effects were observed during the 

experiments. Thus, the outcome of our simulation was to evaluate the accuracy of a readily 

available model to predict the bearing response, which would be desirable to use as a matter of 

convenience. 

Lacking improved models, the parameters of the bilinear model were calibrated independently 

for each simulation. Using the calibrated model, the predicted displacement demand of the 

isolators was within 10% of the observed experimental displacement. When the model was 

calibrated for the peak displacement cycle, it did not capture the history of the displacement 

over smaller cycles very well. Since the demands could not be predicted a priori using this 

approach, an improved simulation model is needed. 

The assumptions used in developing the numerical simulation model that led to a close match 

between the simulated and experimental responses (story drifts, floor accelerations and floor 

spectra) are summarized. The pre-test analytical model incorporated the following features:

• Nonlinear frame elements with composite sections were incorporated to represent the fully 

composite slab-beam interaction, including cyclic demands in the concrete slab. 

• For application of nonlinear frame elements to beams and girders, resultant section behavior 

was modeled in lieu of using fiber sections. The resultant section behavior was calibrated 

based on section analysis of the appropriate fiber section, and the calibrated behavior was 

independent of axial force. 

• Secondary beams were included in the model to represent the vertical vibration 

characteristics. 

• Panel zone springs were included to represent the stiffness of the panel zone. 

Though the structural framing remained essentially in its linear elastic range throughout the test

series, rigorous assumptions improved the accuracy of the numerical simulations. A discussion 

of the influence of these assumptions, which supports the current conclusion, is given in Dao et 

al. (2012b).

Despite the attention to modeling detail, the pre-test model did not predict the structural drifts 

and accelerations well; the primary reason being that the model was overdamped. However, the 
223

following adjustments were made to the post-test simulation model through an iterative trial and 

error process to calibrate the model to the test data.

• The frame elements (beams and girders) were further discretized to about 8 elements per 

structural member to more accurately distribute the mass, which was lumped at the nodes, 

to the system.

• The damping model for the structure was changed from stiffness-proportional damping, 

which damped out the higher mode response, to Rayleigh damping. The damping was 

calibrated to about 2% at the isolation frequency and the higher mode structural frequencies 

observed in the test data. (Note that no viscous damping was applied along the isolation 

system.) An additional damper was added to the model to control the first structural mode. 

The additional damper would not be needed in a program that accommodates constant 

modal damping.

• The isolator vertical stiffness was adjusted by inspection to represent the apparent vertical 

flexibility of the load cell assemblies. Such an adjustment would not be needed for nuclear 

structures.

Following these adjustments, the analytical model matched the floor spectra well; the frequency 

at which the peaks occurred were nearly identical for the analytical model and test data, and 

differences in amplitude occurred due to the lack of precision by which energy dissipation can 

be represented in the structural response. Higher mode participation was non-negligible, 

especially in vertical excitations. 

For nuclear structures, the considerations are:

1. The test results have confirmed that the assumed damping model is probably the single 

most important component for accurate prediction of the structural response. For 

predictability, the modal damping ratios in a representative (e.g. certified design) safety 

related nuclear structure should be evaluated through testing to determine representative 

damping in nuclear structures for future design. The influence of the internal reactor and 

other equipment should be included in such an assessment as nonstructural components 

and content have been shown in the past to influence to markedly influence the damping in 

the system. 

2. The damping in a nuclear structure is expected to be greater than in the tested structure at 

E-Defense. 

3. The large base mat of a nuclear containment structure will filter out some of the higher 

frequency components of the ground motion relative to the free field motions that were 

considered in the test program. This also is expected to reduce higher mode response, 

perhaps significantly, relative to what was observed in the tested structure at E-Defense.

12.3 Hybrid Lead-Rubber (LR) and Cross-Linear (CL) Bearing System

A hybrid isolation system of LR bearings and CL bearings was designed for the test program 

instead of a pure elastomeric isolation system to overcome the constraints of the utilized 

experimental setup. Such constraints would not be applicable in the design of a safety related 

nuclear structure, and thus a hybrid system approach would not be necessary. While a hybrid 

system can overcome stability issues, the vertical force demands in individual bearings are 

difficult to predict due to lack of compliance and subsequent load transfer between the two types 

of devices (LR bearings and CL bearings). Although hybrid systems are used routinely in Japan 
224

for taller structures, the compliance issues suggest that the hybrid system should not be used in 

nuclear facilities. 

The hybrid system was chosen for the following reasons. First, due to the light weight of the 

testbed structure, it was not possible to simultaneously provide the desired period elongation 

and the desired displacement demands with LR bearings alone. Second, the CL bearings 

provided significant tension resistance, which was needed to accommodate the expected tensile 

demands according to preliminary calculations. 

In these tests, the hybrid system resulted in significant axial load transfer between the two types 

of devices; specifically, load redistributed from the LR bearings to the CL bearings as the lateral 

displacement increased, because the rigidity of the base diaphragm constrained the free 

downward movement of the LR bearings. As a positive benefit, the hybrid system eliminated the 

potential that the shear behavior of the LR bearings was affected by stability and post-buckling 

behavior. However, the tradeoff was that the tensile and compressive demands in the CL 

bearings were quite large, as they carried all the overturning induced axial forces, and 

significant tensile demands were observed in the LR bearings, which were constrained by the 

base diaphragm. Prediction of the axial force demands in the individual devices was not 

attempted in our simulation since the forces in the CL bearings were unknown. Inability to 

predict the force demands in design and protect the devices from excessive tension or 

compression is a drawback for the hybrid system.

12.4 Comparison of the Three Building Configurations 

These full scale experiments demonstrated the ability of two different isolation systems to 

significantly attenuate horizontal ground acceleration throughout the structure. Typically, peak 

floor accelerations in the isolated building were 25-50% of the peak input acceleration of the 

simulator platform, while the peak acceleration at the roof level of the fixed-base building under 

the same input motions was amplified by a factor of 2-3 relative to the peak input acceleration. 

This led to a reduction in roof acceleration demand of the isolated buildings by approximately a 

factor of 10 relative to the fixed-base building (Section 11.1).

Different strategies were employed in the design of the two isolation systems. Except for the 

multi-stage behavior of the TP isolation configuration, the design strategies are device 

independent, so that either strategy could be executed with either type of isolation device, 

especially in the absence of mass limitations, which are not a problem for nuclear structures. A 

Moderate Period design strategy was applied to the hybrid LR isolation system, targeting a postyield period of 2.6 sec and a displacement demand of 0.6 meters (23.6 in). A Long Period 

design strategy was applied to the TP isolation system, targeting an “outer slider” period of 5.6 

sec coupled with a displacement capacity exceeding 1 meter (40 in). Although the Long Period 

design allowed for very large displacement demands across the isolators, the benefit of this 

capacity was not fully realized in the test program, as the largest input motion produced a peak 

displacement demand of only 0.7 meter (27.6 in), which was not significantly greater than that 

observed in the Moderate Period design. In other words, the displacement demands for the 

Long Period design did not significantly exceed the displacement demands for the Moderate 

Period design even though some of the motions applied for the Long Period design were 

considered to be much larger intensity. 

The Long Period design led to much larger residual displacements in the isolators than the 

Moderate Period design. The greatest residual displacement observed was about 0.5 cm (0.2 

in) for the Moderate Period design and about 10 cm (4 in) for the Long Period design. (The input 

motions applied to the two systems were not comparable.) In summary, a Long Period design 
225

strategy led to better attenuation of acceleration while not inducing substantially greater 

displacement demands, but at the expense of greater residual displacements. 

Due to the asymmetry of the structural and isolator configuration, base rotation was observed in 

both isolation systems. The torsion in a TP isolation system is theoretically zero since any mass 

asymmetry is naturally balanced by the resisting force in the friction devices, which is a device 

specific effect. However, the “field conditions” introduced a small torsional response since the 

static axial force distribution on the isolators was inconsistent with the distribution of mass in the 

testbed structure. Nevertheless, the base rotation of the TP isolation system was relatively 

small. Greater rotation was observed in the hybrid LR isolation configuration due to the suboptimal configuration of the bearings. This was a problem unique to the restrictive E-Defense 

test setup, and in realistic nuclear structures, base rotations could be eliminated by thoughtful 

placement of isolation devices beneath the base mat. 

12.5 Influence of Vertical Excitation on the Structural Response

Vertical excitation was observed to strongly influence the accelerations observed in the 

structure. Vertical accelerations in columns were amplified slightly relative to the input 

acceleration. The vertical slab accelerations were amplified by a factor of 4-6 compared to the 

input vertical acceleration, which was controlled by the slab vibration properties. The amplitude 

of column and slab vibrations were essentially the same in the hybrid LR isolation configuration 

and the fixed-base configuration. Thus, vertical accelerations in the structure were not amplified 

by the isolation system. The isolation system accommodated the multiple cycles of high 

frequency axial force variation due to vertical excitation. The vertical excitation had a very minor 

influence on the base shear when the intensity of vertical acceleration exceeded 1g (Section 

10.1).

Horizontal floor accelerations were amplified in 3D excitations compared to XY (horizontal only) 

excitations. Amplification of horizontal acceleration occurred in both the isolated and fixed-base 

configurations. Through the analytical investigation, the following source of horizontal-vertical 

coupling was identified. The structural modes of the building, determined by modal analysis of 

the numerical simulation model, were coupled in the horizontal and vertical directions. 

Specifically, vertical floor slab movement was observed in some of the main horizontal structural 

modes of the building, causing those modes to be excited by vertical excitation. The coupling 

was primarily due to a mass irregularity; a large supplemental mass was placed at the roof level 

in an asymmetric configuration in plan. The spectral peaks associated with the coupling modes 

computed from analysis decreased significantly when the supplemental mass in the numerical 

model was moved from the roof and distributed in a regular configuration at the base level. 

Other sources of coupling in the hybrid LR isolation system were hypothesized but not 

confirmed.

The majority of the coupling effects were replicated by a well-crafted numerical simulation model 

that accounted for slab-frame interaction and refined distribution of mass over the floor system. 

The design of the base isolation system and structure should consider and accommodate these 

predictable horizontal-vertical coupling effects.

227

13. REFERENCES

1. Al-Hussaini TM, Constantinou MC, Zayas VA (1994). Seismic Isolation of MultiStory Frame Structures Using Spherical Sliding Isolation System, Technical Report

NCEER-94-0007, State University of New York at Buffalo: Buffalo, NY, USA.

2. American Association of State Highway and Transportation Officials (AASHTO),

(2010). Guide Specification for Seismic Isolation Design, 3rd Edition.

3. American Institute of Steel Construction (AISC), (2005). Steel Construction Manual, 

13rd edition.

4. American Society of Civil Engineers (ASCE), (2008). Seismic Design Criteria for 

Structures, Systems and Components in Nuclear Facilities, ASCE 43-05, ASCE, Reston,

VA.

5. American Society of Civil Engineers (ASCE), (2010). Minimum Design Loads for

Buildings and Other Structures, ASCE/SEI 7-10, ASCE, Reston, VA.

6. Buckle IG, Liu H (1994). “Experimental determination of critical loads of 

elastomeric isolators at high shear strain”, NCEER Bull., 8(3):1–5.

7. Buckle IG, Nagarajaiah S, Ferrell K (2002). “Stability of elastomeric isolation

bearings: Experimental study”, ASCE J. Struct. Eng. 128:3–11.

8. Charney FA, Downs WM (2004). “Modeling procedures for panel zone deformations in 

moment resisting frames”, Proc., ECCS/AISC workshop: Connections in Steel 

Structures V, Amsterdam, Netherlands. 

9. Chiou B, Darragh R, Gregor N, Silva W (2008). “NGA project strong-motion

database”, Earthquake Spectra, 24(1):23–44.

10. Chopra AK (2012). Dynamics of Structures: Theory and Applications to Earthquake 

Engineering, Fourth Edition, Prentice Hall.

11. Clark PW, Aiken ID, Kelly JM (1997). Experimental Studies of the Ultimate Behavior

of Seismically Isolated Structures, Report No. UCB/EERC-97/18, University of 

California: Berkeley, CA, USA.

12. Comite Euro-interational du Beton (CEB), (1996). RC Clements under Cyclic Loading,

State of the Art Report. Thomas Telford Publications, London, England.

13. Constantinou MC, Mokha AS, Reinhorn AM (1990). Experimental and Analytical Study 

of a Combined Sliding Disc Bearing and Helical Steel Spring Isolation System, 

Technical Report NCEER-90-0019, State University of New York at Buffalo: Buffalo, NY,

USA.

14. Constantinou MC, Whittaker AS, Kalpakidis Y, Fenz DM, Warn GP (2007). Performance 

of Seismic Isolation Hardware under Service and Seismic Loading, Technical Report 

MCEER-07-0012, Multidisciplinary Center for Earthquake Engineering Research,

State University of New York at Buffalo: Buffalo, NY, USA.
228

15. Dao ND (2012). Seismic response of a full-scale steel frame building isolated with

triple pendulum bearings under 3D excitation, PhD Dissertation, University of Nevada,

Reno.

16. Dao ND, Ryan KL (2012). “Verification of computational simulation techniques for a 

full- scale steel moment frame building tested at E-Defense”, Submitted to ASCE

Journal of Structural Engineering. 

17. Dao ND, Ryan KL, Sato E, Sasaki T (2013). “Predicting the displacement of triple

pendulum bearings in a full scale shaking experiment using a three-dimensional

element”, Available in Early View, Earthquake Engineering and Structural Dynamics. 

18. Fenz DM, Constantinou MC (2008). Development, Implementation, and Verification

of Dynamic Analysis Models for Multi-spherical Sliding Bearings, Technical Report

MCEER-08-0018, Multidisciplinary Center for Earthquake Engineering Research, State

University of New York at Buffalo, Buffalo, NY, USA.

19. Griffith MC, Kelly JM, Coveney VA, Koh CG (1988a). Experimental Evaluation of

Seismic Isolation of Medium-Rise Structures Subjected to Uplift, Report No. UCB-EERC

88-02, University of California: Berkeley, CA, USA.

20. Griffith MC, Aiken ID, Kelly JM (1988). Experimental Evaluation of Seismic Isolation of

a 9-story Braced Steel Frame Subjected to Uplift, Report No. UCB-EERC 88-05,

University of California: Berkeley, CA, USA.

21. Griffith MC, Aiken ID, Kelly JM (1990). “Displacement control and uplift restraint for 

base-isolated structures”, ASCE J. Struct. Eng., 116:1135–1148.

22. Huang NY, Whittaker AS, Kennedy RP, Mayes RL (2009) “Assessment of BaseIsolated Nuclear Structures for Design and Beyond-Design Basis Earthquake Shaking”, 

Tech. Report MCEER-09-0008, University at Buffalo.

23. Kalpakidis IV, Constantinou MC (2009a). “Effects of heating on the behavior of leadrubber bearings. I: Theory”, J. Struct. Eng. 135:1440–1449.

24. Kalpakidis IV, Constantinou MC (2009b). “Effects of heating on the behavior of leadrubber bearings. II: Verification of theory”, J. Struct. Eng. 135:1450–1461.

25. Kalpakidis IV, Constantinou MC, Whittaker AS (2010). “Modeling strength degradation

in lead-rubber bearings under earthquake shaking”, Earthq. Eng. Struct. Dyn. 39:1533–

1549.

26. Kalpakidis IV, Constantinou MC (2010). “Principles of scaling and similarity for testing

of lead-rubber bearings”, Earthq. Eng. Struct. Dyn. 39:1551–1568.

27. Kasai K, Ooki Y, Ishii M, Ozaki H, Ito H, Motoyui S, Hikino T, Sato E (2008).

“Value- added 5-story steel frame and its components: Part 1 – full scale damper tests

and analyses”, Proc., 14th World Conference on Earthquake Engineering, Beijing, China. 

28. Kasai K, Ito H, Ooki Y, Hikino T, Kajiwara K, Motoyui S, Ozaki H, Ishii M (2010). “Full 

scale shake table tests of 5-story steel building with various dampers,” Proc., 7th

Intern. Conf. on Urban Earthquake Engin. & 5th Intern. Conf. on Earthquake Engin.

Tokyo Inst. Tech., Tokyo, Japan.
229

29. Kasai K, Murata S, Kato F, Hikino T, and Ooki Y (2011). “Evaluation rule for

vibration period, damping, and mode vector of buildings tested by a shake table

with inevitable rocking motions.” J. Struct. Const. Eng., AIJ, 76(270):2031-2040. (In 

Japanese).

30. Kasai K (2011). Personal Communication.

31. Kasalanati A, Constantinou MC (2005). “Testing and modeling of prestressed

isolators”, ASCE J. Struct. Eng. 131:857–866.

32. Kasalanati A (2012). Personal Communication.

33. Kelly JM, Skinner MS, Beucke KE (1980a). Experimental Testing of an Energy

Absorbing Seismic Isolation System, Report No. UCB/EERC-80/35, University of

California, Berkeley: CA, USA.

34. Kelly JM, Beucke KE, Skinner MS (1980b). Experimental Testing of a Friction

Damped Aseismic Base Isolation System with Fail-safe Characteristics, Report No. 

UCB/EERC-80/18, University of California: Berkeley, USA.

35. Kelly JM, Hodder SB (1981). Experimental Study of Lead and Elastomeric Dampers

for Base Isolation Systems, Report No. UCB/EERC-81/16, University of California: 

Berkeley, CA, USA.

36. Kelly JM (1988). Base Isolation in Japan, Report No. UCB/EERC-88/20, University 

of California: Berkeley, CA, USA.

37. Kelly JM, Chalhoub MS (1990). Earthquake Simulator Testing of a Combined

Sliding Bearing and Rubber Bearing Isolation System, Report No. UCB/EERC-87/04, 

University of California: Berkeley, CA, USA.

38. Kelly JM (1997). Earthquake-Resistant Design with Rubber, 2nd Edition, Springer-Verlag. 

39. Kent DC, Park R (1971). Flexural members with confined concrete. Journal of the

Structural Division, Proceedings of the ASCE 97(ST7): 1969-1990.

40. Kikuchi M, Nakamura T, Aiken ID (2010). “Three-dimensional analysis for square

seismic isolation bearings under large shear deformations and high axial loads”, Earthq.

Eng. Struct. Dyn. 39:1513–1531.

41. Koh C-G, Kelly JM (1987). Effects of axial load on elastomeric isolation bearings, Rep.

No. UCB/EERC-86/12, Earthquake Engineering Research Center, Univ. of California,

Berkeley, Calif.

42. Krawinkler H (1978). “Shear in beam-column joints in seismic design of steel 

frames”, Engineering Journal (AISC), 15(3): 82-91.

43. Malushte S and Whittaker AS (2005). "Survey of past base isolation applications in

nuclear power plants and challenges to industry/regulatory acceptance." Proc. 18th

International Conference on Structural Mechanics in Reactor Technology, Beijing, China.

44. McGuire RK, Silva WJ and Costantino CJ (2001). Technical Basis for Revision 

of Regulatory Guidance on Design Ground Motions: Hazard- and Risk-Consistent 
230

Ground Motion Spectra Guidelines, NUREG/CR-6728, U.S. Nuclear Regulatory 

Commission, Washington, D.C.

45. Mokha A, Constantinou MC, Reinhorn AM (1988). Teflon Bearings in Aseismic

Base Isolation: Experimental Studies and Analytical Modeling, Technical Report

NCEER-88-0038, State University of New York at Buffalo: Buffalo, NY, USA.

46. Mokha A, Constantinou MC, Reinhorn AM (1990). Experimental Study and 

Analytical Prediction of Earthquake Response of a Sliding Isolation System with

Spherical Surface, Technical Report NCEER-90-0020, State University of New York at

Buffalo: Buffalo, NY, USA.

47. Monzon EV, Buckle IG, Itani AM (2013). Seismic Performance of Curved Steel Plate

Girder Bridges with Seismic Isolation, Report CCEER 13-06. Center for Civil

Engineering Earthquake Research, University of Nevada, Reno, May, 2013.

48. Morgan TA and Mahin SA (2011). The Use of Innovative Base Isolation Systems to

Achieve Complex Seismic Performance Objectives; Report No. PEER-2011/06,

Pacific Earthquake Engineering Research Center, University of California, Berkeley, 

2011.

49. Nagarajaiah S, Ferrell K (1999). “Stability of elastomeric isolation bearings”, J. Struct.

Eng. 125:946–954.

50. Neuenhofer A, Filippou FC (1997). “Evaluation of nonlinear frame finite-element

models”, Journal of Structural Engineering (ASCE), 123(7): 958-966.

51. Nishiyama I, Okawa I, Fukuyama H, Okuda Y (2011). “Building damage by the 2011 off

the Pacific coast of Tohoku earthquake and coping activities by NILIM and BRI 

collaborated with the administration,” A Report to the U.S.-Japan Cooperative Program in

Natural Resources (UJNR) Panel on Wind and Seismic Effects, Japan Public Works

Research Institute. Retrieved May 15, 2012, From 

http://wwww.kenken.go.jp/english/contents/topics/pdf/report_ujnr2011.pdf

52. Open System for Earthquake Engineering Simulation (OpenSees), Version 2.2.2.

Pacific Earthquake Engineering Research Center, University of California, Berkeley. 

Retrieved May 15, 2010, From http://opensees.berkeley.edu/. 

53. Park YJ, Wen YK, Ang AH-S (1986). “Random vibration of hysteretic systems under 

bi-directional ground motions”, Earthquake Engineering and Structural Dynamics

14:543–557.

54. Ryan KL, Kelly JM, Chopra AK (2005). “Nonlinear model for lead-rubber 

bearings including axial-load effects”, J. Eng. Mech. 131:1270–1278.

55. Ryan K, Sato E, Sasaki T, Okazaki T, Guzman T, Dao N, Soroushian S, Coria C 

(2013a). "Full Scale 5-story Building with Triple Pendulum Bearings at E-Defense", 

Network for Earthquake Engineering Simulation (database), Dataset,

DOI:10.4231/D3X34MR7R

56. Ryan K, Sato E, Sasaki T, Okazaki T, Guzman T, Dao N, Soroushian S, Coria C 

(2013b). "Full Scale 5-story Building with LRB/CLB Isolation System at E-Defense", 

Network for Earthquake Engineering Simulation (database), Dataset,

DOI:10.4231/D3SB3WZ43
231

57. Ryan K, Sato E, Sasaki T, Okazaki T, Guzman T, Dao N, Soroushian S, Coria C 

(2013c). "Full Scale 5-story Building in Fixed-Base Condition at E-Defense", Network for

Earthquake Engineering Simulation (database), Dataset, DOI:10.4231/D3NP1WJ3P

58. Sanchez J, Masroor A, Mosqueda G, Ryan KL (2012). “Static and dynamic

stability of elastomeric bearings for seismic protection of structures”, J. Struct. Eng., In 

Press.

59. Sasaki T, Sato E, Ryan KL, Okazaki T, Mahin SA, and Kajiwara K. “NEES/E-Defense 

base isolation tests: Effectiveness of friction pendulum and lead-rubber bearing

systems”, Proc., 15th World Conference on Earthquake Engineering, Lisbon, 2012.

60. Sato N, Kato A, Fukushima Y, Iizuka M (2002). “Shaking table tests on 

failure characteristics of base isolation system for a DFBR plant”, Nuclear Eng. Des., 

212:293–305.

61. Sato E, Furukawa S, Kakehi A, Nakashima M (2011). “Full-scale shaking table test

for examination of safety and functionality of base-isolated medical facilities”, Earthq.

Eng. Struct. Dyn. 40:1435–1453.

62. Scott BD, Park R, Priestley MJN (1982). “Stress-strain behavior of concrete confined

by overlapping hoops at low and high strain rates”, ACI Journal Proceedings, 79(1): 13-

27.

63. Soroushian S, Ryan KL, Maragakis M, Sato E, Sasaki T, Okazaki T, Tedesco L, Zaghi

AE, Mosqueda G, Alvarez D (2012). “Seismic response of ceiling/sprinkler piping

nonstructural systems in NEES TIPS/NEES nonstructural/NIED collaborative tests on a 

full scale 5-story building”, In Proceedings of the 2012 ASCE Structures Congress, 

Chicago, IL, USA, 29–31 March.

64. Tagawa Y and Kajiwara K (2007). “Controller development for the E-Defense

shaking table”, Proc. IMechE Vol. 221, Part I, J. Systems and Control Engineering. 

65. Takaoka E, Takenaka Y, Nimura A (2011). “Shaking table tests and analysis

method on ultimate behavior of slender base-isolated structure supported by laminated

rubber bearings”, Earthq. Eng. Struct. Dyn. 40:551–570.

66. Uang C-M, Bertero VB (1986). Earthquake Simulation Tests and Associated Studies of a 

0.3- Scale Model of a Six-Story Concentrically Braced Steel Structure, Report No.

UCB/EERC-86/10; University of California: Berkeley, CA, USA.

67. U.S. Code of Federal Regulations (CFR), (2011). Title 10, Part 50, Domestic

Licensing of Production and Utilization Facilities, General Design Criterion (GDC) 2, 

Design Bases for Protection Against Natural Phenomena in Appendix A, General 

Design Criteria for Nuclear Power Plants, U.S. Nuclear Regulatory Commission,

Washington, D.C.

68. Warn GP, Whittaker AS (2006). A Study of the Coupled Horizontal-Vertical Behavior

of Elastomeric and Lead-Rubber Seismic Isolation Bearings, Technical Report 

MCEER-07-0012, Multidisciplinary Center for Earthquake Engineering Research,

State University of New York at Buffalo: Buffalo, NY, USA.

69. Warn GP, Whittaker AS, Constantinou MC (2007). “Vertical stiffness of elastomeric and 

lead- rubber seismic isolation bearings”, ASCE J. Struct. Eng. 133:1227–1236.
232

70. Warn GP, Ryan KL (2012). “A review of seismic isolation for buildings: 

historical developments and research needs”, Buildings (Open Access), In Press. 

71. Weisman J, Warn GP (2012). “Stability of Elastomeric and Lead-Rubber Seismic

Isolation Bearings”, J. Struct. Eng., 138(2), 215–223.

72. Whittaker AS, Uang C-M, Bertero VB (1990). An Experimental Study of the Behavior

of Dual Steel Systems, Report No. UCB/EERC-88/14; University of California: Berkeley, 

CA, USA.

73. Yamamoto S, Kikuchi M, Ueda M, Aiken ID (2009). “A mechanical model for

elastomeric seismic isolation bearings including the influence of axial load”, Earthq. 

Eng. Struct. Dyn. 38:157–180.

74. Zayas V, Low S, Mahin S (1987). The FPS Earthquake Resisting System, Report 

No. UCB/EERC-87/01; University of California: Berkeley, CA, USA.
APPENDIX A 

'(6,1*	&216758&7,21'5$:,1*)257(67,1*2)9$/8($''('

'$03('%8,/',1*

%XLOGLQJ,VRODWHGZLWK+\EULG/HDG5XEEHU,VRODWLRQ6\VWHP

2ULJLQDOVGHYHORSHGE\WKH1DWLRQDO5HVHDUFK,QVWLWXWHIRU(DUWK6FLHQFHDQG'LVDVWHU3UHYHQWLRQ

+\RJR(DUWKTXDNH(QJLQHHULQJ5HVHDUFK&HQWHUIRU9DOXH$GGHG%XLOGLQJ3URMHFW'HFHPEHU

0RGLILHGE\1((67,36SURMHFWIRU1((6('HIHQVH&ROODERUDWLYH7HVW3URJUDPRQ

,QQRYDWLYH,VRODWLRQ6\VWHPV

Architectural Drawings

$ ))3ODQ 6

$D ))3ODQ 6

$ )53ODQ 6

$F (OHYDWLRQ 6

$ (OHYDWLRQ 6

$ 6HFWLRQ 6

$ 'HWDLOHG$UHD 6

$D 'HWDLOHG$UHD 6

$ 6WDLU)ORRU 6

$ 6KDNLQJ7DEOH/D\RXW 6

Structural Drawings 

6WUXFWXUHDQG3DUWLFXODU6SHFLILFDWLRQ

%HDP3ODQV)UDPLQJ(OHYDWLRQ

0DWHULDO&URVV6HFWLRQV

6WHHO-RLQW6WDQGDUG7HVW+RLVWLQJ(TXLSPHQW

)&ROXPQ%DVH

6WHHO6WUXFWXUH

6WHHO6WUXFWXUH

6WXG%ROW/D\RXW

4/'HFN/D\RXW

+LJK'HFN/D\RXW

A-1
A-2
A-3
A-4
A-5
A-6
A-7
A-8
A-9
A-10
A-11
Construction Summary 5 Reinforcement Work 7 Steel Frame Construction 7.6 Welded Joints (Factory Welded Joints)

Value-added Five-story Steel Frame Specimen 5.2 Materials, 5.3 Construction and Assembly 7.1 Common Items 7.6.3

5.2.1 ȷ5.2.2 7.2 Materials

7.2.1 7.6.7

5.3.4

ȷ 7.6.11

6 Concrete Construction6.1 General Notice, 6.2 Concrete Quality

6.1.3 ȷConcrete is as follows:

6.1.4 ȷReceiving Inspection Method by Factory Construction

ȷ Inspection ȷ

6.2.3 7.2.2 ȷ

7.6 Welded Joints (On-site Welded Joints)

6.2.1 a.

Particular Specifications (Building Construction) Structure ȷ

Application of Japanese Architectural Standard Specifications ȷ

1 1) 28 days : standard using water curing or on-site water curing ȷ 7.6.7

2) Over 28 days, less than 91 days : on-site sealed curing

7.2.3

6.2.4 ȷConcrete is as follows

2 ȷ

Documents for Submission

7.2.4

Anchor Bolts

7.10.3

ȷChloride ions of chlorides within the concrete should be 0.3kg/m3 or less.

7.2.7 Deck Plate Type

6.10 Testing6.10.3 ȷReceiving inspections for the concrete used as are follows

6.10.4

7.2.10Material Testing, etc

ȷInspections of the structural concrete are as follows

ȷ

ᲨᲨᲨᲨᲨᲨᲨ

Building ConcreteStrength

Materials andPreparationConditions

* Three test pieces for tensile tests must be made for each lot of steel material used forcolumns, large beams, diaphragms and base plates.* Three test column anchor bolts must be made for the same lot and of the same length.

Flesh ConcreteTestConcrete StrengthTest Provisions6.10.6Building ConcreteStrengthPresumptive Test

Normal BoltInstallation of AnchorBolts

The strength of the structural concrete is shown by the compression strength of test specimens taken from theconstruction site. The difference between the structure concrete strength and test specimen strength is F(=3N/mm2). The curing method of the test specimen has the following ages:

2F to 5F Floor For composite slab Deck plate for composite slabs

Ყ

h=75Ღt=1.2

t=1.2 Flat Deck For frames RF Floor

7 Steel Frame Work Accuracy Measurement Plan and Report58 Reinforced Concrete Work Instructions

Steel Frame Welding Receiving InspectionInstructions and Report

6 Steel Frame Manufacturing Accuracy Measurement Plan and Report3 Steel Frame Construction Factory Manufacturing Instructions4 Steel Frame Construction Factory and On-site Work Plans1 Instructions and Report for each test as shown in the Particular Specifications2 Steel Frame Plans

ȷThe following documents must be submitted to and approved by constructionmanagement before and after the relevant construction.

No Documents for Submission Note

DocumentsforSubmission

Welding (Onsite Welding)

Both the Particular Specification and the “Public Building Construction StandardSpecifications (building construction edition) 2007” (hereinafter referred to as the MinistryStandard Specifications) released by the Administration Division, Government BuildingsDepartment, Minister Secretariat, Ministry of Land, Infrastructure and Transport apply tothe construction of the Vibration Suppression Structure Specimen.Items not listed in the design plan or particular specifications adhere to the MinistryStandard Specifications.Items under other areas of work for each construction adhere to the particularspecifications and the Ministry Standard Specifications (machinery equipment

construction, electrical equipment construction). The chapters, sections and item numbers in this particular specification are linked to therelevant section in the Ministry Standard Specification.

Slip coefficient test must be carried out (Criteria: μ=0.45 or greater)Normal bolts conform to JIS standards or have received certification

QualifiedTechnician(FactoryWelding)Welding(FactoryWelding)QualifiedTechnician(On-siteWelding)

Joined firmly to foundation beam

Steel materials conform to JIS standard products or have received certification,Inspection Certificate to be checked by construction management

Welded SteelJoints and FixingsConcrete TypeConcreteStrengthWorkability andSlump

Joined firmly to steel foundation

National Research Institute for Earth Science and Disaster Prevention HyogoEarthquake Engineering Research Center Design & Construction Drawing forTesting of Value-Added Damped Building

* The attachment base for ALC, curtain wall and measurement apparatus baseare entered into the single item plan.

Reinforcement

Steel MaterialDeck Plate Amount of

Galvanization Type

Qualified welding technicians carrying out factory welding must have passedAW factory welding certification for the welding methods stipulated by AW(Welding position, end tab types)

Input energy, interpass temperature

Steel FrameConstruction SiteHigh-strength Bolt

Use Area Used

The steel frame construction requires approval of construction management.High-strength bolt friction joint surfaces are blast treated

Reinforcement conforms to JIS standards, Inspection Certificate to be checked by construction management.

7.6.3

Note

ConstructionSummarySpecificationsSummary

Welded PartTest

BuildingSummary

Attachment of non-constructed parts

External InspectionVisualUltrasonic Inspection

Instructions for the Receiving Inspection of Complete JointPenetration Welded Parts is created before the test andapproved by construction management.

Qualified welding technicians carrying out factory welding must have passedAW factory welding certification for the welding methods stipulated by AW(Welding position, end tab types)

Input energy, interpass temperature

Compressive Strength

Concrete Temperature

Factory Welding

Sampling Rate

Sampling RateSampling Rate

Same as ministerial certification within the factoryNon-constructed parts should be attached beforehand by factory weldingplates and pieces designed for the non-constructed parts instead of weldingthe parts directly to the constructed material on-site.

Inspection Item Note

Material Material Quality Check material quality on Inspection Certificate

100%

Inspector Construction factory

Conform to "Standards for the Ultrasonic Inspection of WeldDefects in Steel Structures" by the Architectural Institute ofJapan (if there is tensile stress created in the welded areas).

External Inspection

Ultrasonic Inspection

Inspector

Internal

ConstructionFactory Inspectionof Complete JointPenetration WeldedParts

External Inspection,

Ultrasonic Inspection

Factory 100%, On-site 100%

Factory 30%, On-site 100%

CIW certified office

Flux tab (AW-4 certification required)Within ministerial certification

Inspection Item Note

ReceivingInspection ofComplete JointPenetrationWelded Parts

Type Material

Anchor Bolt

(Steel

Foundation)

Double-nut tightening

Material

Ყ

Material Note Type

Normal Bolt SS400 (medium bolt) Self-locking nut

Torshear type S10TųᲢμ=0.45Უ

High-strength bolt JIS type F10TųᲢμ=0.45ᲣHigh-strength bolt

On-site Joint

Column BCR295 Welding Welding

Material Factory Joint Area Used

High-strength bolts conform to JIS standards or have received certification,Inspection Certificate to be checked by construction management

 SS400 Ყ High-strength bolt

Diaphragm Ყ SN490C

Note

High-strength

bolt, welding

 SN490B

Welding

Small beam,IntermediateColumns

Welding ᲧWelding SN490B Large beam

Steel foundation

Slump When test specimens are taken for compression

strength testing, or when quality changes during setting Air Capacity

Chloride Quantity

 When test specimens are taken for compressionstrength testing, or when quality changes during setting When test specimens are taken for compressionstrength testing

Inspection Item Test Item Note

Structural

Concrete

Inspection

Compressive Strength Once for each cast floor, cast section, or cast dayCompression Test Area By a third party organization

Inspection Item

Chloride Quantity When test specimens are taken for compression strength testing

Concrete

Receiving

Inspections

 Once for each cast floor, cast section, or cast day

Slump When test specimens are taken for compression

strength testing, or when quality changes during setting Air Capacity

Air capacity 4.5%

185

Test Item Note

Normal 65 270Normal 65 270 185

Minimum Unit

Water Quantity

kgᲩm3

65 270 185

21

Note

21 24 15 20 (Rubble) 25 (Gravel)

24 15

Lapped

Note

LappedLapped

20 (Rubble)

25 (Gravel)

SD295A

15 20 (Rubble) 25 (Gravel)

SD295A

Maximum Size

Aggregate

mm

Type of

Concrete

SD295A

Normal 21 24

Design

Strength

N/ə2

Quality

Standard

N/ə2

Required

Slump

cm

Area Used Material Properties and Type Joint

Foundation

Construction

Method

Outer wall ALC Plate (t=125, vertical attachment, locking method), glass curtain wall

Inner wall LGS foundation board attachment

Steel

Column Base

Small beam Steel construction

Floor Plate Composite deck plate floor

Roof Steel reinforced concrete slab (frame deck)

Structure Type

Long side Value-added morment steel frame (Y direction)

Short side Value-added morment steel frame (X direction)

Major Parts

and

Material Types

Column Steel construction

Large beam Steel construction

5 floors

Maximum

Height Specimen 1FL+15,835 (RFL slab ceiling)

Bottom of

Foundation Specimen 1FL-900 (Foundation beam bottom)

Construction

Type Steel construction

Revised

Concrete WeightFoundation for Roof Work

Location

1501-21 Nishikameya, Shijimichomitsuda, Miki City, Hyogo PrefectureNational Research Institute for Earth Science and Disaster PreventionHyogo Earthquake Engineering Research Center

Building Use Steel Structured Building Test (Value-added Five-story Steel Frame Specimen)

No. of Floors

Construction Summary, Construction Particular Specifications

12

2008.05.26

S-001

Concrete WeightFoundation forRoof Work

Area Used

Floor SlabFloor SlabConcrete Weight NormalFoundation forRoof Work NormalFloor Slab Normal

Area Used Concrete Type

Maximum

Water/Cement

Ratio

Ჟ

Minimum Unit

Cement Quantity

kgᲩm3

Air capacity 4.5%Air capacity 4.5%

Some supplied by the NationalResearch Institute for Earth Science andDisaster Prevention(S45C: M48 etc)

Type Note

A-12
A-13
A-14
A-15
A-16
A-17
A-18
A-19
A-20
A-21

APPENDIX B 

'5$:,1*6$1'63(&,),&$7,216)25/($'58%%(5%($5,1*6$1'&5266

/,1($5%($5,1*6

%XLOGLQJ,VRODWHGZLWK+\EULG/HDG5XEEHU,VRODWLRQ6\VWHP

&RQWULEXWHGE\'\QDPLF,VRODWLRQ6\VWHPV,QFDQG$VHLVPLF'HYLFHV&R/WG

'HYHORSHGIRU1((6('HIHQVH&ROODERUDWLYH7HVW3URJUDPRQ,QQRYDWLYH,VRODWLRQ6\VWHPV

List of Documents Contributor 

7\SH$,VRODWRU/5%'UDZLQJ '\QDPLF,VRODWLRQ6\VWHPV,QF

,VRODWRU'HVLJQ&DOFXODWLRQV '\QDPLF,VRODWLRQ6\VWHPV,QF

&/%&/%'UDZLQJ $VHLVPLF'HYLFHV&R/WG

&/%6SHFLILFDWLRQ $VHLVPLF'HYLFHV&R/WG

B-1
B-2
EDefense

Isolator Design Calculations

Job No. : Job : EDefense - Final By : AK Date : June 9, 2011

Client : Subject : Seismic Isolator Calculations, LRB Checked : Date : 

Rubber Properties

cs

D Dp

ttp

tip

tbp

N{ ti

Rubber Shear Modulus G  0.060 ksi  G 0.414 MPa 

Rubber's Elongation-at-break H u  5.5

Bearing Dimensions

Overall Diameter D 27.5 in   D 698.5 mm 

Number of Rubber Layers N  40

Lead Diameter Dp  4.0 in Dp 101.6 mm

Shim Thickness ts  0.1196 in ts 3 mm

Layer Thickness ti  0.236 in ti 6 mm

Side Cover Rubber Thickness cs  0.5 in cs 12.7 mm

Top Mounting Plate Thickness ttp  1 in ttp 25.4 mm

Bottom Mounting Plate Thickness tbp  1. in tbp 25.4 mm

Internal Plate Thickness tip  1. in tip 25.4 mm

Isolator Height Hisol N ti  ( )t N 1  s  t  bp t  tp 2 tip   Hisol 18.104 in Hisol 460 mm

800 mm SQ Ext plates; 4 x 1" I Ext holes, 8 x 0.75" I internal connection

Design Displacement

Design Displacement DD  300 mm  Properties are checked at this displacement.

Maximum Displacement DTM  600 mm  Capacity is checked at this displacement.

Project Loads

Rotation appied on the bearing T0  0.0 TDTM  0.0 No rotations are applied on the isolators

Load at undeformed condition Pzero  50 t Assumed

Load at maximum displacement PDTM  50 t

Dynamic Isolation Systems, Inc.

885 Denmark Dr., Suite 101

McCarran, NV 89434

775 359 3333 www.dis-inc.com

LRB Final

Rev. June 9, 2011

B-3
EDefense

Summary

Isolator Dimensions

Overall Diameter D 27.5 in D 698.5 mm 

Number of Rubber Layers N 40 N 40 

Lead Diameter Dp 4 in Dp 101.6 mm

Shim Thickness ts 0.12 in ts 3 mm

Layer Thickness ti 0.236 in ti 6 mm

Side Cover Rubber Thickness cs 0.5 in cs 12.7 mm

Top Mounting Plate Thickness ttp 1 in ttp 25.4 mm

Bottom Mounting Plate Thickness tbp 1 in tbp 25.4 mm

Internal Plate Thickness tip 1 in tip 25.4 mm

Isolator Overall Height Hisol 18.104 in Hisol 459.9 mm

Isolator Properties

600 300 0 300 600

500

250

0

250

500

Design

Displacement

Force

Design Maximum Displacement DD 300 mm

Maximum Corner Displacement DTM 600 mm

Yielded Stiffness Kd 0.65

kN

mm 

Elastic Stiffness Ke 6.5

kN

mm 

Characteristic Strength Qd 65.7 kN

Yield Force Fy 73 kN

Yield Displacement 'y 11.28 mm

Vertical Stiffness Kv 1500

kN

mm 

Shear Force Fmax  DD 259.8 kN Fm'max  DTM 434.6 kN

Effective Stiffness Keff  DD 0.87

kN

mm Km'eff  DTM 0.72

kN

mm 

Energy Dissipated per Cycle EDC D  D 76 kN m EDCm  DTM 155 kN m 

Equivalent Viscous Damping Ratio at Design Displacement E  DD 0.155 Em  DTM 0.094

Shear Strain in Rubber at Design Displacement J  DD 1.25

Shear Strain in Rubber at Maximum Displacement J  DTM 2.5

Allowable Load at Undeformed Condition (with a FS of 3.0) Pallowablezero 4197 kN

Allowable Load at Maximum Displacement PallowableDTM 532 kN

Dynamic Isolation Systems, Inc.

885 Denmark Dr., Suite 101

McCarran, NV 89434

775 359 3333 www.dis-inc.com

LRB Final

Rev. June 9, 2011

B-4
EDefense

Summary

Isolator Dimensions

Overall Diameter D 27.5 in D 698.5 mm 

Number of Rubber Layers N 40 N 40 

Lead Diameter Dp 4 in Dp 101.6 mm

Shim Thickness ts 0.12 in ts 3 mm

Layer Thickness ti 0.236 in ti 6 mm

Side Cover Rubber Thickness cs 0.5 in cs 12.7 mm

Isolator Overall Height Hisol 18.104 in Hisol 459.9 mm

Yielded Stiffness Kd 3.7 kip

in Kd 0.6

kN

mm 

Elastic Stiffness Ke 37 kip

in Ke 6.5

kN

mm 

Characteristic Strength Qd 14.8 kip Qd 65.7 kN

Yield Force Fy 16.4 kip Fy 73 kN

Yield Displacement 'y 0.44 in 'y 11.28 mm

Vertical Stiffness Kv 8566 kip

in Kv 1500

kN

mm 

Undisplaced condition Pallowablezero 4197 kN FS of 3.0

Displacement Minimum of buckling, elastomer limit or a stress limit

DTM  300.mm Pallowable D  TM 4795 kN

DTM  350.mm Pallowable D  TM 4648 kN

DTM  400.mm Pallowable D  TM 3659 kN

DTM  450.mm Pallowable D  TM 2736 kN

DTM  475.mm Pallowable D  TM 2304 kN

DTM  500.mm Pallowable D  TM 1893 kN

DTM  525.mm Pallowable D  TM 1507 kN

DTM  550.mm Pallowable D  TM 1149 kN

DTM  575.mm Pallowable D  TM 822 kN

DTM  600.mm Pallowable D  TM 532 kN

Dynamic Isolation Systems, Inc.

885 Denmark Dr., Suite 101

McCarran, NV 89434

775 359 3333 www.dis-inc.com

LRB Final

Rev. June 9, 2011

B-5
B-6
B-7
B-8
APPENDIX C 

D5$:,1*6)25/2$'&(//$66(0%/,(6$1'&211(&7,2172

7+(6,08/$7253/$7)250

%XLOGLQJ,VRODWHGZLWK+\EULG/HDG5XEEHU,VRODWLRQ6\VWHP

'HYHORSHGE\1((67,363URMHFWIRU1((6('HIHQVH&ROODERUDWLYH7HVW3URJUDPRQ

,QQRYDWLYH,VRODWLRQ6\VWHPV

/HDG&RQWULEXWRU1KDQ''DR

Connection Drawings 

% &RQQHFWLQJ3ODWH3/±(DVW&ROXPQ

% &RQQHFWLQJ3ODWH3/±(DVW&ROXPQ

% &RQQHFWLQJ3ODWH3/±(DVW&ROXPQDQG3ODFHU

% &RQQHFWLQJ3ODWH3/±1RUWK6RXWK:HVW&ROXPQV

% &RQQHFWLQJ3ODWH3/±1RUWK6RXWK:HVW&ROXPQV

% &RQQHFWLQJ3ODWH3/±1RUWK6RXWK:HVW&ROXPQV

% (OHYDWLRQRI/RDG&HOO&RQQHFWLRQ

% &RQQHFWLQJ%HDULQJWR6WUXFWXUH

C-1
C-2
C-3
C-4
C-5
C-6
C-7
C-8
C-9

APPENDIX D 

'5$:,1*6)256758&785$/,167580(17$7,213/$1

%XLOGLQJ,VRODWHGZLWK+\EULG/HDG5XEEHU,VRODWLRQ6\VWHP

'HYHORSHGE\+\RJR(DUWKTXDNH(QJLQHHULQJ5HVHDUFK&HQWHUIRU1((6('HIHQVH&ROODERUDWLYH

7HVW3URJUDPRQ,QQRYDWLYH,VRODWLRQ6\VWHPV

/HDG&RQWULEXWRU7RPRKLUR6DVDNL

Instrumentation Drawings 

6KHHW $FFHOHURPHWHU7DEOH	)

6KHHW $FFHOHURPHWHU)

6KHHW $FFHOHURPHWHU)

6KHHW $FFHOHURPHWHU)

6KHHW $FFHOHURPHWHU)

6KHHW $FFHOHURPHWHU5)

6KHHW $FFHOHURPHWHUVIRU+H[DJRQVKDSHG6WHHO3ODWHV

6KHHW 'LVSODFHPHQW7UDQVGXFHUVIRU%HDULQJV

6KHHW 'LVSODFHPHQW7UDQVGXFHUV)

6KHHW 'LVSODFHPHQW7UDQVGXFHUV)

6KHHW 'LVSODFHPHQW7UDQVGXFHUV)

6KHHW 'LVSODFHPHQW7UDQVGXFHUV)

6KHHW 'LVSODFHPHQW7UDQVGXFHUV)

6KHHW /RDGFHOOVIRU%HDULQJV/5%&/%

6KHHW 6WUDLQ*DJHVRQ&ROXPQ)DFHV)

6KHHW 6WUDLQ*DJHVRQ&ROXPQ)DFHV)

6KHHW 6WUDLQ*DJHVRQ&ROXPQ)DFHV)

6KHHW 6WUDLQ*DJHVRQ&ROXPQ)DFHV)

6KHHW 6WUDLQ*DJHVRQ&ROXPQ)DFHV)

D-1
500 14000 500

2000 16000 2000

Attached on Column

Instrumentation Plan - Accelerometer (Table & 1F)

Accelerometer in X, Y and Z

Accelerometer in vertical direction

SA01 SA02

SA03 SA04

SA05

SA06 SA07

Y

X

SCALE 1:100

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building Tomohiro SASAKI, E-Defense

Sept. 9, 2011

D-2
Instrumentation Plan - Accelerometer (2F)

Attached on Column

Accelerometer in X, Y and Z

Accelerometer in vertical direction, attached at underneath of deck on 2nd floor

SA08

SA09 SA10

SA25

SA26

2500

2500

3500

2500

SA53SA55 SA54

Y

X

SCALE 1:100

Tomohiro SASAKI, E-Defense

Sept. 9, 2011

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building

D-3
Instrumentation Plan - Accelerometer (3F)

Attached on Column

SA11

SA12 SA13

SA27

SA28

2500

2500

3500

2500

Accelerometer in X, Y and Z

Accelerometer in vertical direction, attached at underneath of deck on 3rd floor

SA56

SA58 SA57

Y

X

SCALE 1:100

Tomohiro SASAKI, E-Defense

Sept. 9, 2011

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building

D-4
Instrumentation Plan - Accelerometer (4F)

Attached on Column

SA14

SA15 SA16

SA29

SA30

2500

2500

3500

2500

Accelerometer in X, Y and Z

Accelerometer in vertical direction, attached at underneath of deck on 4th floor

SA59

SA61 SA60

Y

X

SCALE 1:100

NEES-E-Defense Collaborative Project - Shake Table Experiment on Base-Isolated Steel Building Tomohiro SASAKI, E-Defense

Sept. 9, 2011

D-5
Instrumentation Plan - Accelerometer (5F)

Attached on Column

SA17

SA18 SA19

SA31

SA32

2500

2500

3500

2500

Accelerometer in X, Y and Z

Accelerometer in vertical direction, attached at underneath of deck on 5th floor

SA62

SA64 SA63

SA66

Accelerometer in vertical direction, attached at top surface of deck on 5th floor

Y

X

SCALE 1:100

Tomohiro SASAKI, E-Defense

Sept. 9, 2011

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building

D-6
Instrumentation Plan - Accelerometer (RF)

Accelerometer in X, Y and Z

Accelerometer in vertical direction

SA20

SA21 SA22

SA33

SA34

2500

2500

3500

2500

SA65

SA67 SA68

3500

SCALE 1:100

Tomohiro SASAKI, E-Defense

Y

X

Sept. 9, 2011

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building

D-7
Instrumentation Plan - Accelerometers for Hexagon-shaped Steel Plates

Accelerometer (Horizontal)

SA45X

SA45Y

SA47X

SA47Y

SA49X

SA49Y

SA51X

SA51Y

Y

X

SCALE 1:100

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building Tomohiro SASAKI, E-Defense

Sept. 9, 2011

D-8
2.5m

2.5m

2.5m

1.2m 1.2m 1.2m

WirePod

Direction of wire

Target

Instrumentation Plan - Displacement Transducers for Bearings

SD01 SD02 SD03

SD04SD05SD06

Y

X

SCALE 1:100

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building Tomohiro SASAKI, E-Defense

Sept. 9, 2011

D-9
Lasor DT

Direction of lasor

Target

Instrumentation Plan - Displacement Transducers (1F)

SD23 SD24

SD26 SD25

190

460

460

150

Y

X

SCALE 1:100

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building Tomohiro SASAKI, E-Defense

Sept. 9, 2011

D-10
Instrumentation Plan - Displacement Transducers (2F)

Lasor DT

Direction of lasor

Target

SD27

SD28

SD29

SD30

2050

2310

2502.5

2050

Y

X

SCALE 1:100

Tomohiro SASAKI, E-Defense

Sept. 9, 2011

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building

D-11
Instrumentation Plan - Displacement Transducers (3F)

Direction of lasor

Target

SD31

SD32

SD34

SD33

2060

2295

2050

2502.5

Lasor DT

Y

X

SCALE 1:100

Tomohiro SASAKI, E-Defense

Sept. 9, 2011

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building

D-12
Instrumentation Plan - Displacement Transducers (4F)

Lasor DT

Direction of lasor

Target

SD35

SD36

SD38

SD37

2050

2300

2050

2497.5

Y

X

SCALE 1:100

NEES-E-Defense Collaborative Project - Shake Table Experiment on Base-Isolated Steel Building Tomohiro SASAKI, E-Defense

Sept. 9, 2011

D-13
Instrumentation Plan - Displacement Transducers (5F)

Lasor DT

Direction of lasor

Target

SD39

SD40

SD42

SD41

2050

2300

2050

2495

Y

X

SCALE 1:100

Tomohiro SASAKI, E-Defense

Sept. 9, 2011

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building

D-14
Loadcell A (xy250kN/z400kN)

Loadcell B (xy400kN/z700kN)

Instrumentation Plan - Loadcells for Bearings (LRB/CLB)

SL10 SL11

SL12

SL13

SL14

SL15

SL16

SL42 SL43

SL44

SL45

SL46

SL47

SL48

SL33 SL34

SL35

SL36

SL37

SL38

SL39

SL40

SL41

SL58 SL59

SL60

SL61

SL62

SL63

SL64

Y

X

SCALE 1:100

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building Tomohiro SASAKI, E-Defense

Sept. 9, 2011

D-15
SS04L: 1580 above from top surface of base beam

SS04U: 670 below from bottom of beam on 2nd floor

Instrumentation Plan - Strain Gauges on Column Faces (1F)

Strain gauge

1 3

4

2

Y

X

SCALE 1:100

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building Tomohiro SASAKI, E-Defense

Sept. 9, 2011

D-16
SS08L: 505 above from top surface of deck on 2nd floor

SS08U: 670 below from bottom of beam on 3rd floor

Strain gauge

Instrumentation Plan - Strain Gauges on Column Faces (2F)

1 3

4

2

Y

X

SCALE 1:100

Tomohiro SASAKI, E-Defense

Sept. 9, 2011

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building

D-17
Strain gauge

Instrumentation Plan - Strain Gauges on Column Faces (3F)

SS12L: 505 above from top surface of deck on 3rd floorSS12U: 670 below from bottom of beam on 4th floor

1 3

4

2

Y

X

SCALE 1:100

Tomohiro SASAKI, E-Defense

Sept. 9, 2011

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building

D-18
Strain gauge

Instrumentation Plan - Strain Gauges on Column Faces (4F)

SS16L: 505 above from top surface of deck on 4th floorSS16U: 670 below from bottom of beam on 5th floor

1 3

4

2

Y

X

SCALE 1:100

NEES-E-Defense Collaborative Project - Shake Table Experiment on Base-Isolated Steel Building Tomohiro SASAKI, E-Defense

Sept. 9, 2011

D-19
Strain gauge

Instrumentation Plan - Strain Gauges on Column Faces (5F)

SS20L: 505 above from top surface of deck on 5th floorSS20U: 670 below from bottom of beam on roof floor

1 3

4

2

Y

X

SCALE 1:100

Tomohiro SASAKI, E-Defense

Sept. 9, 2011

E-Defense/NEES Collaborative Research Project - Shake Table Experiment on Base-Isolated Steel Building

D-20
APPENDIX E 

,62/$7257(675(3257)25/($'58%%(5%($5,1*6

%XLOGLQJ,VRODWHGZLWK+\EULG/HDG5XEEHU,VRODWLRQ6\VWHP

&RQWULEXWHGE\'\QDPLF,VRODWLRQ6\VWHPV,QF

'HYHORSHGIRU1((6('HIHQVH&ROODERUDWLYH7HVW3URJUDPRQ

,QQRYDWLYH,VRODWLRQ6\VWHPV

E-1
Dynamic Isolation Systems, Inc.

885 Denmark Drive, Suite 101  McCarran, NV 89434 USA

Tel:1Ͳ775Ͳ359Ͳ3333  Fax:1Ͳ775Ͳ359Ͳ3985

www.disͲinc.com

ISOLATOR TEST REPORT

PROJECT NAME: NRC

DATE ISSUED: JULY 11, 2011

REPORT NUMBER: 152Ͳ10

PREPARED BY: DIS INC.

E-2
Introduction

This report has been prepared for Isolator Types A and B for the NRC project.Included in this report are

the isolator testing plans, testing results, drawings, an overview of the testing facility, and the test rig

calibration certificates.

The testing was performed on July 7, 2011 at DIS' test facility in McCarran, NV.

Table 1ͲTested Bearing Serial Numbers

Bearing Type Serial Number

A 16439

16443

B 16450

16458

E-3
Table of Contents 

LIST OF TABLES .......................................................................................................................... II

LIST OF FIGURES ......................................................................................................................... II

TEST MATRIX TABLE ................................................................................................................. 1

TEST MATRIX .............................................................................................................................. 1

COMPRESSION SHEAR TESTS ...................................................................................................... 1

MEASUREMENT & CALCULATION PROCEDURES ................................................................. 4

COMPRESSION SHEAR TESTS ...................................................................................................... 4

TEST RESULTS AND HYSTERESIS LOOPS ............................................................................. 5

DISCUSSION OF TEST RESULTS ................................................................................................... 5

SUMMARY OF TEST RESULTS ...................................................................................................... 6

HYSTERESIS LOOPS .................................................................................................................... 7

TEST APPARATUS ................................................................................................................... 16

OVERVIEW OF TEST FACILITY .................................................................................................... 16

TEST RIG CALIBRATION ............................................................................................................. 17

List of Tables 

Table 1 - Tested Bearing Serial Numbers ...................................................................................... i

Table 2 - Isolator Type A - Test A Results .................................................................................... 6

Table 3 - Isolator Type A - Test B Results .................................................................................... 6

Table 4 - Isolator Type A - Test C Results .................................................................................... 6

Table 5 - Isolator Type A - Test D Results .................................................................................... 6

List of Figures 

Figure 1 - Combined Compression and Shear Test Matrix ........................................................... 2

Figure 2 - Isolator Type A Drawing ............................................................................................... 3

Figure 3 - Graphical Display of Isolator Shear Properties ............................................................. 4

Figure 4 - DIS Big Rig Photo ....................................................................................................... 16

Figure 5 - DIS Big Rig Drawing ................................................................................................... 16

E-4
Test Matrix Table

Test Matrix

The test procedure is determined by a compilation of test information called "Test Matrix".  The

contents of the test matrix, such as the order of testing and number of tests, are determined by the

designer (or the parties involved in the project).The test matrices for this project are presented in this

section.

Compression Shear Tests

Each test is described by one row of information in the test matrix table.Each test performed bears a

test ID, shown in the first column of the test matrix table.

The testing parameters are:

a) Compression load on the bearing to be maintained during the test.

b) Displacement deformation to be imposed on the isolator.

c) Number of fullyͲreversedͲcycles of testing.

d) The acceptance criteria for the test, if any.

One cycle of shear deformation consists of movement from the zero position to the maximum specified

displacement in the positive direction (Dmax), then to the maximum specified displacement in the

negative direction (Dmin), and back to the zero position.  This movement is applied in a smooth

continuous way, similar to a constant velocity sawͲtooth shape.

E-5
Combined Compresssion and Shear Test Approval Rev Date Description

KF 0 06/29/11 Issued for Approval

Isolator Type A

A 1 1.7 600 3 125 300 Note 1B 2 2.8 1000 3 208 500 Note 1

C 3 0.3 100 0.5 271 650 Note 1

D 4 1.7 600 3 125 300 Note 1

Notes:1. Wait 15 minutes between tests

Revision Status

Test 

ID Notes Test Order C. Stress (N/mm2) Shear Strain (%) Displ. (mm) Load (kN) Number of Cycles

Des. Eng. 

Approval NRC Proj. Mng. 

Approval Production Isolators

Test Sup. 

Approval

Combined Compression & Shear

Test Matrix No. 152-100-01

Customer 

Approval

E-6
E-7
Measurement & Calculation Procedures

Compression Shear Tests

Testing begins when the axial load is applied to the isolators.Maintaining the axial load, the isolators

are sheared to the specified displacement for the appropriate number of cycles. Using the recorded

shear force and displacements from the test, shear forceͲdisplacement plots (hysteresis loops) are

generated.Since two isolators of the same type are tested simultaneously, the total measured shear

force has been multiplied by one half during processing to produce hysteresis loops for a single isolator.

The essential properties of an isolator can be extracted from a hysteresis loop.  Figure 2 shows an

example of a hysteresis loop generated after a compression shear test and some of the isolator

properties that are obtained from it.Three properties are measured directly from the recorded data,

the maximum isolator displacement (Dmax), the maximum force required to displace the isolator (Fmax),

and the area of the hysteresis loop which gives the total energy dissipated per cycle (EDC).The effective

stiffness of the isolator (Keff) is equal to Fmax/Dmax.  There is no engineering judgment or estimation

involved the the determination of Fmax, Dmax, Keff, or EDC.Fmax and Dmax are selfͲevident and EDC is

determined by numerical integration of the recorded forceͲdisplacement data file.

The hysteresis loop has the following properties in addition to the measured properties listed above.

These are the hysteretic force at zero displacement (Qd), the yielded stiffness of the isolator (K2), and the

unloading stiffness of the isolator (K1).A biͲlinear loop is fitted to the actual hysteresis loop such that

the fitted loop has identical measured properties (Keff and EDC) as the actual hysteresis loop.

Figure 2ͲGraphical Display of Isolator Shear Properties

E-8
Test Results and Hysteresis Loops

Discussion of Test Results

The compression shear and compression stiffness test results are summarized in Table 2 and Error!

Reference source not found. as well the corresponding hysteresis loops.The hysteresis loops exhibit

positive incremental stiffness and the isolators remained stable during testing.

E-9
Summary of Test Results 

Table 2ͲIsolator Type AͲTest A Results

Serial

Numbers

Axial Load Dmax Kd Qd F0 Keff

(kN) (mm) (kN/cm) (kN) (kN) (kN/cm)

16439 600 300 6.3 41.4 64.8 8.4

16443 600 300 6.3 41.4 64.8 8.4

16450 600 300 6.2 39.2 64.5 8.3

16458 600 300 6.2 39.2 64.5 8.3

Average 600 300 6.3 40.3 64.7 8.4

Table 3ͲIsolator Type AͲTest B Results

Serial

Numbers

Axial Load Dmax Kd Qd F0 Keff

(kN) (mm) (kN/cm) (kN) (kN) (kN/cm)

16439 1000 500 5.1 45.9 75.7 6.6

16443 1000 500 5.1 45.9 75.7 6.6

16450 1000 500 5.0 43.6 75.5 6.5

16458 1000 500 5.0 43.6 75.5 6.5

Average 1000 500 5.1 44.8 75.6 6.6

Table 4ͲIsolator Type AͲTest C Results

Serial

Numbers

Axial Load Dmax Kd Qd F0 Keff

(kN) (mm) (kN/cm) (kN) (kN) (kN/cm)

16439 100 650 5.5 56.2 85.1 6.8

16443 100 650 5.5 56.2 85.1 6.8

16450 100 650 5.4 53.1 86.1 6.8

16458 100 650 5.4 53.1 86.1 6.8

Average 100 650 5.5 54.7 85.6 6.8

Table 5ͲIsolator Type AͲTest D Results

Serial

Numbers

Axial Load Dmax Kd Qd F0 Keff

(kN) (mm) (kN/cm) (kN) (kN) (kN/cm)

16439 600 300 5.8 41.7 63.4 7.9

16443 600 300 5.8 41.7 63.4 7.9

16450 600 300 5.8 40.1 62.8 7.9

16458 600 300 5.8 40.1 62.8 7.9

Average 600 300 5.8 40.9 63.1 7.9

E-10
Hysteresis Loops

E-11
Force (kN)

Displacement (cm)

-30 -20 -10 10 20 30 

 60 

-60

 120 

-120

 180 

-180

 240 

-240

Job: 152 (NRC) Test Name : 25838-001.dat

Class: Production Type: A Isolators 16439 & 16443 Tested: 7/7/2011 5:55:42 PM

Test Type: Production Test Matrix ID: A (Stress: 1.7N/mm^2, Strain: 125%)

Cycle Dmax(cm) Fmax(kN) Keff(kN/cm) Qd(kN) EDC(kN.cm) K2fit(kN/cm) Fofit(kN) V(cm/min)

 1 30.12 266.15 8.84 45.46 8454.7 6.47 71.24 136.09

 2 30.15 248.61 8.24 40.03 7486.1 6.16 62.90 140.41

 3 30.21 245.26 8.12 38.61 7206.2 6.12 60.40 140.41

AVERAGE 30.16 253.34 8.40 41.4 7715.7 6.25 64.84 139.0

E-12
Force (kN)

Displacement (cm)

-30 -20 -10 10 20 30 

 60 

-60

 120 

-120

 180 

-180

 240 

-240

Job: 152 (NRC) Test Name : 25834-001.dat

Class: Production Type: A Isolators 16450 & 16458 Tested: 7/7/2011 2:28:21 PM

Test Type: Production Test Matrix ID: A (Stress: 1.7N/mm^2, Strain: 125%)

Cycle Dmax(cm) Fmax(kN) Keff(kN/cm) Qd(kN) EDC(kN.cm) K2fit(kN/cm) Fofit(kN) V(cm/min)

 1 30.17 264.55 8.77 43.29 8505.8 6.40 71.49 136.33

 2 30.16 245.63 8.14 37.70 7433.0 6.08 62.38 140.40

 3 30.16 240.83 7.99 36.49 7108.8 6.01 59.63 140.46

AVERAGE 30.16 250.33 8.30 39.2 7682.5 6.16 64.49 139.1

E-13
Force (kN)

Displacement (cm)

-50 -40 -30 -20 -10 10 20 30 40 50 

 100 

-100

 200 

-200

 300 

-300

Job: 152 (NRC) Test Name : 25839-001.dat

Class: Production Type: A Isolators 16439 & 16443 Tested: 7/7/2011 6:27:49 PM

Test Type: Production Test Matrix ID: B (Stress: 2.8N/mm^2, Strain: 208%)

Cycle Dmax(cm) Fmax(kN) Keff(kN/cm) Qd(kN) EDC(kN.cm) K2fit(kN/cm) Fofit(kN) V(cm/min)

 1 50.25 344.95 6.86 47.36 16181.4 5.25 81.19 138.12

 2 50.21 329.41 6.56 45.74 14809.4 5.08 74.31 140.96

 3 50.21 325.42 6.48 44.73 14284.9 5.05 71.65 140.94

AVERAGE 50.23 333.26 6.64 45.9 15091.9 5.13 75.72 140.0

E-14
Force (kN)

Displacement (cm)

-50 -40 -30 -20 -10 10 20 30 40 50 

 100 

-100

 200 

-200

 300 

-300

Job: 152 (NRC) Test Name : 25835-001.dat

Class: Production Type: A Isolators 16450 & 16458 Tested: 7/7/2011 2:54:06 PM

Test Type: Production Test Matrix ID: B (Stress: 2.8N/mm^2, Strain: 208%)

Cycle Dmax(cm) Fmax(kN) Keff(kN/cm) Qd(kN) EDC(kN.cm) K2fit(kN/cm) Fofit(kN) V(cm/min)

 1 50.22 336.42 6.70 44.52 16198.4 5.08 81.26 138.25

 2 50.20 320.95 6.39 43.31 14690.8 4.93 73.67 141.01

 3 50.21 316.35 6.30 42.97 14278.0 4.88 71.57 141.03

AVERAGE 50.21 324.57 6.46 43.6 15055.7 4.96 75.50 140.1

E-15
)RUFHN1

'LVSODFHPHQWFP

           

















-RE15&7HVW1DPHDGDW

&ODVV3URGXFWLRQ7\SH',VRODWRUV	7HVWHG$0

7HVW7\SH3URGXFWLRQ7HVW0DWUL[,'&6WUHVV1PPA6WUDLQ

+&\F 'PD[FP )PD[N1 .HIIN1FP 4GN1 ('&N1FP .ILWN1FP )RILWN1 9FPPLQ

        

E-16
)RUFHN1

'LVSODFHPHQWFP

           

















-RE15&7HVW1DPHDGDW

&ODVV3URGXFWLRQ7\SH',VRODWRUV	7HVWHG$0

7HVW7\SH3URGXFWLRQ7HVW0DWUL[,'&6WUHVV1PPA6WUDLQ

+&\F 'PD[FP )PD[N1 .HIIN1FP 4GN1 ('&N1FP .ILWN1FP )RILWN1 9FPPLQ

        

E-17
Force (kN)

Displacement (cm)

-30 -20 -10 10 20 30 

 60 

-60

 120 

-120

 180 

-180

 240 

-240

Job: 152 (NRC) Test Name : 25841-001.dat

Class: Production Type: A Isolators 16439 & 16443 Tested: 7/7/2011 7:16:44 PM

Test Type: Production Test Matrix ID: D (Stress: 1.7N/mm^2, Strain: 125%)

Cycle Dmax(cm) Fmax(kN) Keff(kN/cm) Qd(kN) EDC(kN.cm) K2fit(kN/cm) Fofit(kN) V(cm/min)

 1 30.13 247.63 8.22 44.42 7954.8 5.99 67.20 135.77

 2 30.14 235.37 7.81 41.13 7439.2 5.73 62.74 140.41

 3 30.14 232.66 7.72 39.48 7146.6 5.72 60.23 140.43

AVERAGE 30.14 238.55 7.92 41.7 7513.6 5.81 63.39 138.9

E-18
Force (kN)

Displacement (cm)

-30 -20 -10 10 20 30 

 60 

-60

 120 

-120

 180 

-180

 240 

-240

Job: 152 (NRC) Test Name : 25837-001.dat

Class: Production Type: A Isolators 16450 & 16458 Tested: 7/7/2011 3:38:45 PM

Test Type: Production Test Matrix ID: D (Stress: 1.7N/mm^2, Strain: 125%)

Cycle Dmax(cm) Fmax(kN) Keff(kN/cm) Qd(kN) EDC(kN.cm) K2fit(kN/cm) Fofit(kN) V(cm/min)

 1 30.15 246.41 8.17 42.47 7938.8 5.95 66.97 135.85

 2 30.17 233.82 7.75 39.54 7350.9 5.70 61.89 140.58

 3 30.10 230.44 7.65 38.24 7058.4 5.68 59.51 140.56

AVERAGE 30.14 236.89 7.86 40.1 7449.3 5.78 62.79 139.0

E-19
Test Apparatus

Overview of Test Facility

The DIS test facility is located at its manufacturing plant in McCarran, NV.The testing laboratory has

approximately 4,000 square feet of available floor space and houses two combined compression and

shear test rigs along with their support hardware and is serviced by a 10Ͳton overhead crane.The big

test rig is illustrated in Figure 3 and Figure 4 below.As shown in the figures, two isolators are tested

together in the hydraulically powered test rig.The isolators can be examined for both their shear and

compression properties in the test rig.

Figure 3ͲDIS Big Rig Photo

Figure 4ͲDIS Big Rig Drawing

E-20
Test Rig Calibration

E-21
E-22
E-23
E-24
E-25







NUREG/CR-7196 Large Scale Earthquake Simulation of a Hybrid Lead Rubber Isolation System 

Designed with Consideration of Nuclear Seismicity