vertical earthquake loads on seismic isolation systems in bridges

Vertical Earthquake Loads on Seismic IsolationSystems in Bridges

Gordon P. Warn1 and Andrew S. Whittaker, M.ASCE2

Abstract: An important consideration for the design of seismic isolation systems composed of elastomeric and lead–rubber bearings isthe safety of individual bearings for maximum considered earthquake shaking. One assessment of bearing safety involves the calculationof the vertical �or axial� earthquake load on the individual seismic isolation bearings. This paper investigates the influence of verticalearthquake excitation on the response of a bridge isolated with low-damping rubber and lead–rubber bearings through earthquakesimulation testing. Response data collected from the experimental program are used to determine the vertical load on the isolation systemdue to the vertical component of excitation. A comparison of the normalized vertical load data to the vertical base acceleration showedsignificant amplification of the vertical response for each simulation and configuration. Disaggregation of the axial load history showedthe summation of maximum values from the vertical earthquake load and overturning moment overestimates the maximum axial loadbecause these maximum values are unlikely to occur simultaneously. Additionally, a spectral analysis procedure using the unreducedvertical stiffness of the bearings was shown to provide a reasonable estimate of the vertical earthquake load.

DOI: 10.1061/�ASCE�0733-9445�2008�134:11�1696�

CE Database subject headings: Isolation; Earthquakes; Simulation; Bridges; Seismic effects.

Introduction

Seismic isolation is a method for reducing inertial forces thatdevelop in a structure as a result of earthquake ground shaking.Inertial forces are reduced by lengthening the fundamental periodof vibration and added damping through the introduction of ele-ments �isolators� with low horizontal and large vertical stiffnessthat decouple the superstructure from the supporting substructure.Elastomeric bearings are one type of isolator consisting of a num-ber of elastomeric �rubber� layers bonded to intermediate steel�shim� plates. The horizontal flexibility �low shear stiffness� of anelastomeric bearing is dictated by the total thickness of rubber,whereas the close spacing of the intermediate shim plates pro-vides a large vertical �relative to the shear� stiffness for a givenbonded rubber area and elastomer shear modulus. Lead–rubberbearings are another type of seismic isolation hardware and arederived from the elastomeric bearing differing only through theaddition of a lead core �an energy dissipating mechanism� typi-cally located in a central hole.

A key aspect of the design of seismic isolation systems forbridges �and buildings� is the stability of individual bearings dur-ing maximum considered earthquake shaking. Once maximumdisplacements across the isolation interface are determined, the

1Assistant Professor, Dept. of Civil and Environmental Engineering,Penn State Univ., University Park, PA 16802. E-mail: [email protected]

2Professor, Dept. of Civil, Structural and Environmental Engineering,State Univ. of New York at Buffalo, Buffalo, NY 14260.

Note. Associate Editor: Marvin W. Halling. Discussion open untilApril 1, 2009. Separate discussions must be submitted for individualpapers. The manuscript for this paper was submitted for review and pos-sible publication on December 14, 2006; approved on March 21, 2008.This paper is part of the Journal of Structural Engineering, Vol. 134,No. 11, November 1, 2008. ©ASCE, ISSN 0733-9445/2008/11-1696–
1704/$25.00.
1696 / JOURNAL OF STRUCTURAL ENGINEERING © ASCE / NOVEMBER

safety of the individual bearings is assessed. The safety checkrequires the calculation of: �1� the reduced critical buckling loadof the elastomeric bearing at the maximum horizontal displace-ment and �2� the compressive load under maximum earthquakeshaking consisting of dead load, live load, overturning �if appli-cable�, and vertical earthquake components. Determination of theaxial load carrying capacity of an elastomeric �and lead–rubber�bearing under lateral displacement is an important topic and onethat has been investigated by others, e.g., Buckle et al. �1994,2002�. Previous analytical work �Button et al. 2002� investigatingthe effect of vertical ground motion on the seismic response of avariety of highway bridges. The results of this study illustratedthe significant effect the vertical component of excitation canhave on pier axial load and vertical deck shear forces. Although,three of the six bridge systems considered in the Button et al.study included elastomeric bearings �modeled as linear springs�for the response-spectrum and linear response-history analysis,the primary focus was the response of the bridge deck and piers.The work presented in this paper focuses on axial load demand onthe seismic isolation systems and individual bearings, or morespecifically, the contribution of axial load due to the vertical com-ponent of excitation.

This paper summarizes the results of an earthquake simulationprogram performed with a large scale bridge model isolated withlow-damping rubber �LDR� and lead-rubber �LR� bearings aimedat investigating the influence of vertical ground motion on theload carried by the isolation system and the individual bearings.Axial load histories from the individual bearings recorded duringearthquake simulation testing were utilized to investigate the con-tribution of the overturning and vertical earthquake componentsto the total axial load carried by the individual bearings. In addi-tion, sample results from selected earthquake simulation are usedto validate a spectral analysis procedure for estimating verticalearthquake loads on seismic isolation system and individual bear-
ings utilizing an unreduced vertical stiffness.
2008

Bridge Model and Bearings

The influence of vertical excitation on the load carried by a seis-mic isolation system and individual bearings was studied using asingle span steel truss superstructure supported by four seismicisolation bearings. A photograph of the isolated bridge model in-stalled on the two earthquake simulators housed in the StructuralEngineering and Earthquake Simulation Laboratory �SEESL� atthe University at Buffalo, Buffalo, N.Y. is presented in Fig. 1.Reduced-scale LDR and LR seismic isolation bearings were de-signed and fabricated for component and earthquake simulationtesting.

The bearings were proportioned assuming mechanical and ma-terial properties typical of prototype elastomeric and lead–rubberbearings used in bridge construction �HITEC 1998a,b�, a bilinearforce displacement relationship: identical to the characterizationassumed in the Guide Specifications for Seismic Isolation Design�AASHTO 1999� and scaling procedures for dynamic testing�Harris and Sabnis 1999�. The basis for the design is the assumedtarget prototype period of vibration, related to the shear stiffnessof the rubber and the supported weight, of 2.5 s that was relatedto the model period using a time scale factor �St� of 2. Additionalprototype bearings properties assumed for the design include: astatic compressive pressure �p� of 3.45 MPa �500 psi�; a charac-teristic strength normalized by the weight acting on the isolator�Qd /W� of 0.09 �LR only�; and effective damping ratios, at adisplacement equivalent to 100% rubber shear strain, of �5 and20% for the LDR and LR bearings, respectively.

Details of the as-built LDR and LR bearings are presented inFig. 2. As illustrated in Fig. 2, the LDR and LR bearings are

Fig. 1. Photograph of isolated bridge model

Fig. 2. Details of LDR and LR bearings

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identical with the exception of a lead-core inserted in the centralhole for the LR bearings. In plan, the bearings consist of a hollowcircular cross section with an outer bonded rubber diameter �do�of 152 mm and an inner diameter of 30 mm. In elevation, thebearings consist of 20 individual rubber layers each 3 mm thick�0.125 in.� �total rubber thickness �Tr� of 60 mm �2.36 in.��bonded to 19 intermediate steel shim plates each 3 mm�0.125 in.� thick and two 25 mm �1.0 in.� thick internal endplates. The specified outer bonded rubber diameter was reducedfrom 172 to 152 mm for manufacturing purposes resulting inshape factors �S� of 10.2 and 12.2 for the LDR and LR bearings,respectively, where the shape factor is defined as the ratio of theloaded area to the area free to bulge for a single layer. For the LRbearing, the lead-core prevents the rubber along the inner radiusfrom bulging.

The bridge superstructure consists of two parallel planartrusses spanning 10.7 m �35 ft� each 1.5 m �5 ft� in height�center-to-center� and transversely spaced 1.2 m �4 ft� apart�center-to-center�. In the transverse direction, the planar trussesare connected by lateral beams and diagonal cross bracing. Theself-weight of the truss superstructure is 89 kN �20 kips�. Threeadditional mass packages were added to the model each consist-ing of 2 steel plates �3,000�2,000�90 mm�, 120 lead bricks,and some ancillary steel section together weighing, approxi-mately, 90 kN and resulting in a total model weight of, approxi-mately, 360 kN. Due to the reduced bonded rubber diameter�152 mm� and existing experimental setup, the provided staticpressure in the isolated configuration was 5.2 MPa �750 psi� perbearing and judged to be within the representative range of designcompressive pressure for prototype bridge bearings �HITEC1998a,b�. Table 1 presents frequency and critical damping ratiovalues for the first vertical mode of the loaded truss-bridge in thefixed-base �FB� and isolated �LDR and LR� configurations. Forthe fixed-base configuration, the vertical frequency and dampingratio of the first vertical mode were determined from the results ofwhite-noise testing to be 12 Hz �0.0833 s� and 2%, respectively�calculated using the half-power bandwidth method�. For the iso-lated configurations �LDR and LR� the vertical frequency andcritical damping ratio were estimated using the results of compo-nent testing in conjunction with the results of white-noise testingin the FB configuration. The estimated vertical frequency andcritical damping ratio for the LDR configuration were estimatedto be 10.2 Hz and 1%, respectively. Similarly, for the LR configu-ration the vertical frequency and critical damping ratio were esti-mated to be 10.8 Hz and 2%, respectively.

Component Testing

Two LDR and two LR bearings, identical to those used for earth-quake simulation testing, were dedicated to component testingthat consisted of characterization tests, to determine the mechanic

Table 1. Dynamic Properties of Fixed-Base and Isolated Bridge

Systemf

�Hz��

�%�

Fixed-basea 12 2

LDRb 10.2 1

LRb 10.8 2aValues determined from the results of white-noise testing.bValues estimated from results of component testing.

properties of the bearings, and lateral-offset tests, to investigate

F STRUCTURAL ENGINEERING © ASCE / NOVEMBER 2008 / 1697

the influence of lateral displacement on the vertical stiffness. Re-sults of the lateral-offset program are presented in Warn andWhittaker �2006� and Warn et al. �2007�. In this section, sampleresults from characterization testing performed on the LDR andLR bearings are presented. Testing was carried out using thesingle bearing testing machine located in the SEESL at the Uni-versity at Buffalo. A photograph of the testing machine is pre-sented in Fig. 3. A five-channel reaction load cell �identical tothose used for earthquake simulation testing� located directly be-neath the seismic isolation bearing was used to measure shear andaxial forces and a linear variable displacement transducer housedin the horizontal actuator used to measure lateral displacement�see Fig. 3�. For tests with variable axial loading, two linear po-tentiometers were attached across the bearing end plates and usedto measure relative vertical displacement.

Fig. 4 presents sample horizontal force–displacement loops forthe LDR �Fig. 4�a�� and LR �Fig. 4�b�� bearings from a test con-sisting of four fully reversed cycles of sinusoidal input to a dis-placement amplitude of 60 mm �100% rubber shear strain� at afrequency of 0.01 Hz with a constant compressive pressure of3.45 MPa. From the results presented in Fig. 4�a�, the effectivestiffness �Keff� and effective damping ratio ��eff� of the LDR,determined from the third-cycle response, are 0.29 kN /mm

Fig. 3. Photograph of single bearing testing machine

−80 −60 −40 −20 0 20 40 60 80−40

−30

−20

−10

0

10

20

30

40

Displacement (mm)

She

arfo

rce

(kN

)

(a)

−80 −60 −40 −20 0 20 40 60 80−40

−30

−20

−10

0

10

20

30

40

Displacement (mm)

She

arfo

rce

(kN

)

(b)

Fig. 4. Horizontal force–displacement response of bearings: �a�LDR; �b� LR


�1.7 kips / in.� and 2.8%, respectively. Additionally, the effectiveshear modulus �Geff� was estimated to be 0.83 MPa calculated forthe LDR bearings as

Keff =GeffAe

Tr�1�

where Ae=effective area accounting for the contribution of the12 mm thick cover �see Fig. 2� determined to be equal to 1.2times the bonded rubber diameter �Ab�; see Warn and Whittaker�2006� for details. The effective stiffness �Keff� and effectivedamping ratio ��eff� of the LR �see Fig. 4�b��, again determinedfrom the third-cycle response, are 0.4 kN /mm �2.3 kips / in.� and19%, respectively. In addition, the effective yield strength ��l� ofthe lead core was estimated to be 8.4 MPa �1,200 psi� and calcu-lated according to the following relationship:

Qd = �lAl �2�

where Al=cross-sectional area of the lead core and Qd

=characteristic strength. The test described previously was re-peated at a frequency of 1 Hz to determine the rate dependency ofthe mechanical properties of the bearings. From the results of the1 Hz test, a 7% increase in both Keff and �eff was observed for theLDR bearings. For the LR bearings, Keff and �eff increased by 25and 10%, respectively. The substantial increase in both Keff and�eff for the LR bearings is attributed, in part, to the increase ineffective yield strength of the lead core at the 1 Hz frequency.

Sample force–displacement results from axial load tests per-formed on the LDR and LR bearings are presented in Fig. 5,wherein, a positive value of axial load indicates compression. Theaxial load tests were conducted under force control with a rampsignal, subjecting each bearing to three cycles of load at 0.01 Hz:from zero to the maximum and back to zero. Due to the mass ofthe loading beam and the stability and safety of the machine inforce control, the axial load tests could not be conducted at afrequency greater than 0.33 Hz. However, from a 0.33 Hz test

0 0.5 1 1.5 20

50

100

150

200

Vertical displacement (mm)

Axi

allo

ad(k

N)

(a)

0 0.5 1 1.5 20

50

100

150

200

Vertical displacement (mm)

Axi

allo

ad(k

N)

(b)

Fig. 5. Vertical force–displacement response of bearings: �a� LDR;�b� LR

repeated on the LDR and LR bearings, only a slight ��10% �

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increase in vertical stiffness was observed �Warn and Whittaker2006�. From the results presented in Fig. 5�a�, the vertical stiff-ness �Kv0� of the LDR bearing, calculated as the secant stiffnesson the ascending branch of the first cycle, was determined to be84 kN /mm �480 kips / in.� and the vertical effective damping ratio��v� was estimated to be 1.2%. From the results presented in Fig.5�b�, Kv0 and �v for the LR bearing were determined to be163 kN /mm and 2%, respectively. Although not the intended pur-pose, the addition of the lead core increases the vertical stiffness,e.g., from 84 kN /mm �LDR� to 163 kN /mm �LR� and to a lesserextent the effective vertical damping, �v, from 1.2% �LDR� to,approximately, 2% �LR�.

Earthquake Ground Motions

Four sets of recorded earthquake ground motions were selectedand obtained from the Pacific Earthquake Engineering Researchdatabase �2005� for the purpose of earthquake simulation testing.Table 2 presents summary information for the selected earthquakeground motion records including: mechanism; moment magni-tude; station; distance-to-fault; and site class/classification. Eachset consists of two horizontal components and one vertical com-ponent. Two of the sets were recorded in close proximity to afault �Sylmar and Japan Meteorological Agency �JMA��, one ofwhich �Sylmar� possesses characteristics of forward-rupture di-rectivity �Somerville 2000�. Elastic response spectra generatedassuming 5% of critical damping are plotted in Fig. 6. Also plot-ted in Fig. 6 is the ratio of the vertical response �denoted UP� toeach horizontal response: V /H. For the Hyogo-Ken �Fig. 6�b��and Sylmar �Fig. 6�d�� motions, recorded in close proximity to afault, the spectral ratio exceeds the commonly assumed value of2 /3 �indicated by the solid horizontal line� for periods less than0.25 s, and for one component of the Hyogo-Ken motion for pe-riods between 1.0 and 2.5 s. However, for the Bolu �Fig. 7�a�� andRio Dell �Fig. 7�c�� motions, recorded at approximately 20 km tothe fault, the spectral ratio is approximately less than or equal to2 /3. The trends of the V /H spectral ratio presented here are con-sistent with those observed by other researchers for near- andfar-field ground motion records �Silvia 1997; Button et al. 2002�.

For the purpose of earthquake-simulation testing all groundmotion records were time scaled by a factor of 0.5 in accordancewith the scaling procedure used to proportion the bearings andamplitude scaled according the earthquake simulation testing pro-gram described in the next section.

Earthquake Simulation Testing

Earthquake simulation testing was performed on the two 6degree-of-freedom earthquake simulators housed in the SEESL,serving the George E. Brown Jr., Network for Earthquake Engi-

Table 2. Summary of Earthquake Ground Motion Records

Year Country Event Mechanism

1992 United States Cape Mendocino Reverse normal

1994 United States Northridge Reverse normal

1995 Japan Hyogo-Ken Nanbu Strike slip

1999 Turkey Duzce Strike slipaDistance value represents closest to fault rupture.

neering Simulation Equipment site, at the University at Buffalo

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�Fig. 1�. Four different configurations were tests, specifically, twoisolation systems �composed of LDR or LR� and two differenttransverse support configurations. Illustrations of the transverseview of the isolated bridge model with the bearings in the 1.8 and1.2 m support width configuration are presented in Fig. 7. Foreach ground motion set, earthquake simulations were performedat two intensity levels: �1� either 25 or 50% to verify the perfor-mance of the earthquake simulators, the instrumentation and tocompare measured bearing properties to those used for analyticalpredictions and �2� 50–100%, as representative design groundshaking.

Several instruments were used to measure and record theresponse of the isolated bridge model and input motion of theearthquake simulators including, accelerometers, string potenti-ometers, load cells and a Krypton Portable Coordinate Trackingand Measurement System. The accelerometers and string potenti-ometers were distributed across the isolated bridge model to mea-sure and record absolute accelerations and displacements. Inaddition, a five-channel reaction load cell was located beneatheach seismic isolation bearing recording axial load, shear force inthe x �longitudinal� and y �transverse� directions, and momentsabout the x and y axes; see Fig. 7.

A key response quantity for this study is the vertical load onthe isolation system �Peq� due to the vertical component of exci-tation also referred to herein as the vertical earthquake load. Forsimulations performed with three components of excitation �hori-zontal plus vertical�, the vertical earthquake load, Peq, was calcu-lated using the recorded normal �or axial� load signals from eachof the four load cells according to

Peq�ti� = �j=1

4

Pj�ti� − Wt �3�

where Pj�ti�=axial load signal from the jth load cell at time ti andWt=total static weight of the bridge model calculated as the sumof the initial readings from each load cell. In Eq. �3� the contri-bution of the overturning moment to the axial loads is removed inthe summation.

The vertical earthquake load, Peq, was normalized by the gen-eralized weight of the bridge model �W*� to investigate the am-plification of the vertical response and to calculate amplificationfactors. The generalized weight of the bridge model was esti-mated to be 204 kN �46 kips� and calculated by approximatingthe superstructures as a lumped mass system with a parabolicdeformed shape corresponding to a shape function of �n

= �0.48,1 ,0.48�T at 1,524, 5,334 �mid-span�, and 9,144 mm alongthe length of the bridge model. To illustrate the amplification ofthe vertical response, the Peq /W* and vertical base accelerationhistories from a simulation with the LR bearing in the 1.8 mconfiguration and the JMA ground motion set conducted at 100%intensity are presented in Fig. 8. Base input acceleration recordedat the center of the east and west earthquake simulator extension

entitude Station

Distancea

�km� Site class/classification

1 Rio Dell Overpass 18.5 B/USGS

7 Sylmar–Olive View 6.4 C/USGS

9 JMA 0.6 B/USGS

1 Bolu 17.6 C/USGS

Mommagn

7.

6.

6.

7.

platforms are presented in Figs. 8�b and c�, respectively. Noting,


some difference is observed between the maximum input accel-eration from the east �maximum=0.37 g� and west �maximum=0.28 g�. From the results presented in Fig. 8, significant ampli-fication is observed in the vertical response �1.4 g� despite thedifference in peak input from the east �0.37 g� or west �0.28 g�platforms. The significant amplification of the vertical response isa result of the flexibility of the isolation-truss system �Warn andWhittaker 2006� and the low level of damping provided by theisolation system ��2% of critical for the LR and LDR� and thesteel truss ��2% of critical�. Amplification factors �� were cal-culated from the results of each simulation as

0 0.5 1 1.5 2 2.5 3 3.50

0.5

1

1.5

2

2.5

3

T (s)

Sa

(g)

Bolu 0Bolu 90Bolu UP

(a)

0 0.5 1 1.5 2 2.5 3 3.50

0.5

1

1.5

2

2.5

3

T (s)

Sa

(g)

JMA 0JMA 90JMA UP

(c)

0 0.5 1 1.5 2 2.5 3 3.50

0.5

1

1.5

2

2.5

3

T (s)

Sa

(g)

Rio 270Rio 360Rio UP

(e)

0 0.5 1 1.5 2 2.5 3 3.50

0.5

1

1.5

2

2.5

3

T (s)

Sa

(g)

Sylmar 90Sylmar 360Sylmar UP

(g)

Fig. 6. Elastic response spectra and spectral ratios for unscaled motion


� =max�Peq/W*�

PSA�4�

where PSA�average peak simulator �base� acceleration calcu-lated from the maximum absolute value of the recorded accelera-tion from the east and west simulation platforms. Fig. 9 presentsamplification factors calculated from the results of simulationswith the LDR and LR isolation system in both the 1.8 and 1.2 mconfigurations. In Fig. 9, the simulation from which the amplifi-cation value was calculated is identified by an abbreviation for theground motion set and the intensity level. The � results presented

0 0.5 1 1.5 2 2.5 3 3.5 40

0.5

1

1.5

2

T (s)

V/H

Rat

io

UP/90UP/360

)

0 0.5 1 1.5 2 2.5 3 3.5 40

0.5

1

1.5

2

T (s)

V/H

Rat

io

UP/90UP/360

)

0 0.5 1 1.5 2 2.5 3 3.5 40

0.5

1

1.5

2

T (s)

V/H

Rat

ioUP/90UP/360

0 0.5 1 1.5 2 2.5 3 3.5 40

0.5

1

1.5

2

T (s)

V/H

Rat

io

UP/90UP/360

)

5% critical damping: �a� Bolu; �b� JMA; �c� Rio Dell; and �d� Sylmar

4(b

4(d

4(f)

4(h

s and

in Fig. 9 illustrate the significant amplification in the vertical re-

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sponse for each isolation system, support width configuration andearthquake ground motion record with values ranging from, ap-proximately 2 �LR 1.2, Sylmar 75%� to 5.5 �LDR 1.2, JMA 50%�.

The axial load history on the individual bearings was disag-gregated to investigate the contribution of the individual compo-nents �i.e., overturning and vertical excitation� to the totalresponse. The component of axial load due to overturning wascalculated by subtracting the initial recorded static weight and aquarter of the vertical load calculated according to Eq. �3� fromthe response recorded by a particular load cell. Fig. 10 presentsaxial load histories for LR 1 from a simulation performed with theJMA ground motion set at 100% intensity. In Fig. 10, the staticweight acting on LR 1 �97 kN� was subtracted from the recordedaxial load prior to disaggregating the signal for comparison. FromFig. 10�a�, the maximum and minimum values of the overturningcomponent are 47 and −54 kN, respectively, and from Fig. 10�b�,

Fig. 7. Transverse view of isolated bridge

0 5 10 15 20−1.5

−1

−0.5

0

0.5

1

1.5

Peq

/W*

max.=1.4min.=−1.3

(a)

0 5 10 15 20−1.5

−1

−0.5

0

0.5

1

1.5

Acc

eler

atio

n(g

) max.=0.37min.=−0.32

(b)

0 5 10 15 20−1.5

−1

−0.5

0

0.5

1

1.5

Time (s)

Acc

eler

atio

n(g

) max.=0.28min.=−0.27

(c)

Fig. 8. �a� Normalized vertical load and vertical input accelerationsfrom the �b� east and �c� west supports for the LR system with JMAat 100% intensity

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the maximum and minimum values from the vertical componentare 95 and −95 kN, respectively. As expected, the frequency ofthe vertical component �Fig. 10�b�� is significantly higher than theoverturning response: governed by the horizontal response of theisolators. From the axial load history presented in Fig. 10�c��overturning plus vertical�, the maximum and minimum valuesare 110 and −106 kN, respectively. A key observation from Fig.10�c�, is that the maximum absolute value �110 kN� is signifi-cantly less than the sum of the maximum absolute values of theoverturning �Fig. 10�a�� and axial �Fig. 10�b�� load �149 kN�.This result suggests that it would be overly conservative to addthe maximum vertical component to the maximum overturningcomponent for the assessment of stability during a maximum con-sidered event as these are unlikely to occur simultaneously, al-though the maximum axial load due to overturning is certain tooccur at the maximum horizontal displacement.

Spectral Analysis

In this section the vertical seismic force on the isolation system isestimated using a spectral analysis procedure whereby the bridgeand isolation system are treated as an equivalent linear systemand the response is estimated from spectra generated using therecorded seismic input and an appropriate level of damping. It isimportant to note that the vertical response of elastomeric andlead–rubber bearings is highly nonlinear especially under tensileloading and high compressive pressures. However in compressionand at compressive pressures that would be considered acceptable

l and support width: �a� 1.8 m; �b� 1.2 m

LDR 1.2 LDR 1.8 LR 1.2 LR 1.80123456789

10

Rio

Del

l100

%

Bol

u50

%

JMA

50%

Rio

Del

l100

%

Bol

u50

% JMA

50%

Bol

u10

0%

JMA

100%

Syl

mar

75%

Bol

u10

0%

JMA

100%

Syl

mar

75%

Am

plifi

catio

n

Fig. 9. Amplification factors for vertical motion from earthquakesimulation testing

mode


for design, elastomeric and lead–rubber bearings exhibit reason-ably linear behavior as illustrated by the vertical force–displacement response of the LDR and LR bearings presented inFig. 5.

To investigate the impact of the reduction in vertical stiffnessexhibited by elastomeric and lead–rubber seismic isolators �Warnand Whittaker 2006; Warn et al. 2007� the spectral analysis wasperformed twice using bounding values of the vertical stiffness ofthe seismic isolator, specifically, the vertical stiffness under zerolateral offset �Kv0� and a reduced vertical stiffness �Kv� account-ing for the maximum isolator displacement. In the subsequentsection, the vertical seismic forces calculated using the spectrummethod are compared to the vertical earthquake load �Peq� deter-mined from the results of earthquake simulation testing to evalu-ate the accuracy of the simplified method, understanding that fordesign purposes the seismic hazard �input� is not deterministic.The vertical seismic force on the isolation system was calculatedas

Pv = SaW* �5�

where Sa=spectral acceleration and W*=effective weight as pre-viously defined. For this calculation, Sa was determined from theaverage of the response spectra generated using the recorded seis-mic input from the east �ACC9� and west �ACC3� simulator ex-tension platforms. Fig. 11 presents elastic response spectragenerated using recorded acceleration histories from a simulationwith the LR bearings at 1.2 m support width and the JMA groundmotion set at 100% intensity and 2% critical damping. Theequivalent period of the bridge-isolation system, assuming full

0 5 10 15 20−150

−100

−50

0

50

100

150

Axi

allo

ad(k

N)

(a)

0 5 10 15 20−150

−100

−50

0

50

100

150

Axi

allo

ad(k

N)

(b)

0 5 10 15 20−150

−100

−50

0

50

100

150

Time (s)

Axi

allo

ad(k

N)

(c)

Fig. 10. Axial load histories for LR 1 at 1.8 m support width andJMA 100% intensity: �a� overturning; �b� vertical; and �c� overturningplus vertical

vertical stiffness �Kv0� was calculated as


Tv0 = 2� W*

Keq,v0 · g�6�

where g=gravitational acceleration constant, and Keq,v0 is theequivalent stiffness of the bridge-isolation system calculated as

1

Keq,v0=

1

Kbv+

1

4Kv0�7�

where Kbv=generalized vertical stiffness of the truss-bridge deter-mined �from the results of white-noise tests performed in thefixed-base configuration� to be approximately 120 kN /mm andKv0=unreduced �zero lateral displacement� vertical stiffness ofeither an individual LDR or LR bearing. The effective period ofthe bridge-isolation system �Tv� using the reduced vertical stiff-ness �Kv� of the LDR and LR bearings was calculated in a similarmanner, however, the lower bound value Kv was estimated ac-cording to

Kv =Kv0

�1 +12

2

R2� �8�

where =maximum horizontal displacement of the isolator �de-termined from the results of earthquake simulation testing� andR=outer radius of the bearing. The expression presented in Eq.�8� was derived from the Koh–Kelly two-spring model �Koh andKelly 1987; Warn et al. 2007�.

Comparison of Results

Values of the vertical earthquake load calculated using the spec-tral analysis �Pv� procedure were compared to the vertical loadresults determined from earthquake simulation �Peq� to evaluatethe accuracy of the spectral analysis procedure.

Plotted in Fig. 12 is a comparison of the estimated verticalload �Pv� with the experimentally determined vertical load �Peq�from simulation performed with the LDR �Fig. 12�a�� and LR�Fig. 12�b�� isolation systems. Each plot includes three referencelines with slopes equal to 1.0 �solid�, 0.85 �dashed�, and 1.15�dashed�. Again, the vertical earthquake load �Pv� was calculatedconsidering the full vertical stiffness of the isolators under zerolateral displacement �Kv0� and again with a reduced vertical stiff-ness �Kv� as denoted in the legend of the plots. From the resultspresented in Fig. 12, the vertical earthquake load calculated usingthe spectral analysis procedure and the full vertical stiffness isobserved to estimate the experimentally determined value �Peq�

0 0.25 0.5 0.75 10

0.5

1

1.5

2

T (s)

Sa

(g)

ACC3ACC9Avg.

Fig. 11. Elastic response spectra for 2% critical damping withrecorded acceleration histories: LR 1.2 JMA at 100% intensity

with reasonable accuracy: most values lying in close proximity of

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�15% of Peq for both the LDR �Fig. 12�a�� and LR �Fig. 12�b��isolation systems. Additionally, for this isolation system andbridge model, consideration of the reduced vertical stiffness in thespectral analysis calculation did not result in improved estimatesof the vertical earthquake load, rather, use of the reduced verticalstiffness lead to more unconservative estimates �Pv� Peq�.

To assess the contribution of the isolator to the total verticalamplification a comparison of amplification factors for the FB andisolated configurations �LDR-bridge and LR-bridge� determinedfrom spectral analysis �denoted SA� is presented in Fig. 13. Alsoplotted are experimentally determined amplification factors forthe LDR- and LR-Bridge in the 1.2 m configuration. From thespectral analysis �SA� results presented in Fig. 13, only a mar-ginal increase in amplification is predicted for the isolated con-figurations over the fixed-base configuration with the exception ofthe JMA motion. However amplification factors for the isolatedconfigurations calculated using SA tend to underestimate �to vary-ing degrees� the experimentally determined values with theexception of the LR-bridge and the Sylmar ground motion.

Discussion and Conclusions

This paper serves to summarize and present sample results from

0 50 100 150 200 250 300 350 400 4500

50

100

150

200

250

300

350

400

450

Peq

(kN)

Pv

(kN

)

Kvo

1.8 mK

vo1.2 m

Kv

1.8 mK

v1.2 m

(a)

0 50 100 150 200 250 300 350 400 4500

50

100

150

200

250

300

350

400

450

Peq

(kN)

Pv

(kN

)

Kvo

1.8 mK

vo1.2 m

Kv

1.8 mK

v1.2 m

(b)

Fig. 12. Comparison of vertical earthquake load: �a� LDR isolationsystem; �b� LR isolation system

earthquake simulation testing performed on a bridge model iso-

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lated with LDR and LR bearings. Results from the earthquakesimulation testing program were used to investigate the influenceof vertical excitation on the vertical load carried by the isolationsystem and the axial load of individual bearings. In addition, theresults of earthquake simulation testing were used to evaluate aspectral analysis procedure for calculating the vertical earthquakeload on the isolation system considering the full and a reducedvertical stiffness of the LDR and LR bearings.

Significant amplifications in the vertical response for both theLDR- and LR-bridge configurations were experimentally ob-served. However from a comparison of amplification factors forboth the isolated and fixed-base configurations estimated usingspectral analysis suggests the isolation system itself results inonly a marginal increase in amplification over the fixed-basebridge for the model and systems considered in this study. Theseresults suggest that the vertical flexibility of the bridge-isolationsystem should not be ignored for design and that use of the peakground acceleration of the vertical component �y-axis ordinate ofthe design spectrum� would underestimate the vertical earthquakeload on the isolation system.

Sample axial load histories from individual bearings showedthe frequency of the vertical component to be significantly higherthan that of the overturning component as was expected. How-ever, the sum of the maximum absolute value of the two compo-nents tends to overestimate the maximum absolute value of thecombined �overturning plus vertical� axial load history for allsimulations performed in this study. The spectral analysis proce-dure considering the full vertical stiffness of the isolator lead toreasonably accurate estimates of the vertical earthquake load onthe isolation system for this bridge model and isolation systems.However, for a hybrid isolation system such as those composed offlat sliding and elastomeric the reduction in vertical stiffnessshould be considered on a case-by-case basis using more ad-

Rio Dell 100% Bolu 50% JMA 50%0

1

2

3

4

5

6

Am

plifi

catio

n

SA: FBSA: LDR 1.2EXP: LDR 1.2

(a)

Bolu 100% JMA 100% Sylmar 75%0

1

2

3

4

5

6

Am

plifi

catio

n

SA: FBSA: LR 1.2EXP: LR 1.2

(b)

Fig. 13. Comparison of amplification for fixed-base �FB� with: �a�LDR–bridge; �b� LR–bridge systems

vanced analysis techniques.


Acknowledgments

The writers gratefully acknowledge the financial support of theMultidisciplinary Center for Earthquake Engineering Researchand the Federal Highway Administration through Task D1.5 ofthe Federal Highway Administration Contract No. DTFH 61-98-C-0094. The writers also wish to thank Dr. Amarnath Kasalanatiof DIS Inc. for providing the model bearings used in this study.The opinions expressed in this paper are those of the writers anddo not reflect the opinions of the Multidisciplinary Center forEarthquake Engineering Research or the Federal Highway Ad-ministration. No guarantee regarding the results, findings, andrecommendations are offered by either the Multidisciplinary Cen-ter for Earthquake Engineering Research or the Federal HighwayAdministration.

References

AASHTO. �1999�. Guide specifications for seismic isolation design,Washington, D.C.

Buckle, I., Nagarajaiah, S., and Ferrell, K. �2002�. “Stability of elasto-meric isolation bearings: Experimental study.” J. Struct. Eng., 128�1�,3–11.

Buckle, I. G., and Liu, H. �1994�. “Experimental determination of criticalloads of elastomeric isolator at high shear strain.” NCEER Bull., 8�3�,1–5.

Button, M. R., Cronin, C., and Mayes, R. L. �2002�. “Effect of verticalmotions on the seismic response of highway bridges.” J. Struct. Eng.,128�12�, 1551–1564.


Harris, H. G., and Sabnis, G. M. �1999�. Structural modeling and experi-mental techniques, CRC, Boca Raton, Fla.

HITEC. �1998a�. Evaluation findings for Dynamic Isolation Systems, Inc.elastomeric bearings, Civil Engineering Research Foundation, Wash-ington, D.C.

HITEC. �1998b�. Evaluation findings for Skellerup base isolation elasto-meric bearings, Civil Engineering Research Foundation, Washington,D.C.

Koh, C. G., and Kelly, J. M. �1987�. “Effects of axial load on elastomericisolation bearings, Springfield, Va., Berkeley, Calif.” Earthquake En-gineering Research Center, College of Engineering, Univ. of Califor-nia, Berkeley, Calif.

Pacific Earthquake Engineering Research �PEER�. �2005�, “Strong mo-tion database.” �http://peer.berkeley.edu/� �May 1, 2005�.

Silva, W. J. �1997�. “Characteristics of vertical ground motions for appli-cations to engineering design.” Proc., FHWA/NCEER Workshop onthe National Representation of Seismic Ground Motion for New andExisting Highway Facilities, Tech. Rep. No. NCEER-97–0010, Na-tional Center for Earthquake Engineering Research, State Univ. ofNew York at Buffalo, N.Y., 205–252.

Somerville, P. �2000�. “Characterization of near-fault ground motions.”U.S.–Japan Workshop on the Effects of Near-Field Earthquake Shak-ing, Pacific Earthquake Engineering Research Center, San Francisco.

Warn, G., and Whittaker, A. S. �2006�. “A study of the coupledhorizontal-vertical behavior of elastomeric and lead-rubber seismicisolation bearings.” Tech. Rep. No. MCEER-06-0010, Multidisci-plinary Center for Earthquake Engineering Research, State Univ. ofNew York at Buffalo, Buffalo, N.Y.

Warn, G. P., Whittaker, A. S., and Constantinou, M. C. �2007�. “Verticalstiffness of elastomeric and lead–rubber seismic isolation bearings.” J.Struct. Eng., 133�9�, 1227–1236.

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vertical earthquake loads on seismic isolation systems in bridges

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