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SEISMIC PROTECTION OF LIGHT SECONDARY SYSTEMSTHROUGH DIFFERENT BASE ISOLATION SYSTEMSMAURO DOLCE a & DONATELLO CARDONE aa DiSGG - Department of Structural Engineering , University of Basilicata Contrada Macchia Romano ,85100, Potenza, ItalyPublished online: 03 Jun 2008.

To cite this article: MAURO DOLCE & DONATELLO CARDONE (2003) SEISMIC PROTECTION OF LIGHT SECONDARY SYSTEMS THROUGHDIFFERENT BASE ISOLATION SYSTEMS, Journal of Earthquake Engineering, 7:2, 223-250, DOI: 10.1080/13632460309350447

To link to this article: http://dx.doi.org/10.1080/13632460309350447

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Journal of Earthquake Engineering, Vol. 7, NO. 2 (2003) 223-250 @ lrnperid College Press @ Imperial College Press www.icprtr$.co.uk

SEISMIC PROTECTION OF LIGHT SECONDARY SYSTEMS THROUGH DIFFERENT BASE

ISOLATION SYSTEMS

MAURO DOLCE' and DONATELLO CARDONE+ DiSGG - Department of Structud Engineering,

University oJ Basilicata, Contmda Macchia Romann,

85100, Potenzn, Italy *doIcemmOlibem. it

tdonntelio. cOtiscalinet.it

Received 14 January 2001 Revised 23 June 2002

Accepted 26 June 2002

The objective of the prgent work is to m i n e advantages and drawbacks of different types of isolation systems, when seismic isolation is used as a protection strategy against damage to internal equipment and contents. The starting point of the study is the big experimental program of table tests on reduced-scale R/C structural models, carried out within the MANSIDE (Memory Alloys for New Seismic Isolation DEvices) project. Seven identical 133-scaled, bstorey frames were tested, including two fixed-base mod- els and four base-isolated models with different isolation systems, namely: (1) rubber isolators, (2) steel-hysteretic system and (3), re-centring SMA (Shape Memory Alloy) system. In this study the internal equipment is regarded as an elastic single degree of freedom, with 2% equivalent viscous damping. Therefore, the capability of fixed-base and base-isolated models with different isolation systems to protect light secondary sys- tems is evaluated by comparing the floor response spectra obtained from the storey accelerations recorded during shaking table tests. Three different PGA's are considered, about 0.15g, 0.39 and 0.59, respectively. All the shaking table tests are also simulated with an accurate numerical mnodel, to validate and better understand the experimental results. It is found that each type of isolation system reduces considerably the seis mic effects on internal equipments in wide frequency regions. However, tuning effects may arise in specific frequency ranges, corresponding to ,the first mode in structures equipped with quasi-elastic (rubber) isolation systems, and to higher modes in struc- tures equipped with elast-plastic (steel) and nonlinear re-centering (SMA) isolation systems.

Keywortfa: Base isolation; rubber bearings; shape memory alloys; internal equipment; shaking table tests; floor response spectra.

1. Introduction

In many cases, the monetary value of the non structural damage caused by an earthquake greatly exceeds the value of structural damage. Non structural damage includes both damage to non structural members (such as infill masonry panels, ceilings, windows, etc.) and equipment (electrical and mechanical systems), as well

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224 M. Dolce Ed D. Cardone

as to internal goods (furniture, tanks, statues or sophisticated equipments). The two types of non structural damage are mainly governed by interstorey drifts and by floor accelerations, respectively. For buildings having strategic importance in post-earthquake emergency (e.g. hospitals, fireman stations, etc.) or hgh-risk con- tent (e.g. thermal or nuclear power plants, etc.) or high-value content (e.g. muse- ums, monuments, computer centres, etc.), the need to protect non structural parts and internal content under strong earthquakes is so strong as to affect structural design.

Base isolation [Naeim and Kelly, 19991 seems to be the most effective sekimic protection strategy against non structural damage, due to the strong reduction of both interstorey drifts and floor accelerations. Nevertheless, it presents some limitations. As a matter of fact, elastic or quasi-elastic isolation systems tends to concentrate all3hXinputiSmic energy in a quite narrow spectral range, in which the natural frequency of objects or equipment can fall [Suarez and Singh, 19871. .---o- n-thEth- -- .,- - -- . ____._..--._-- - ~ -

er hand strongly nonhear isolation systems are usually considered lit- tle effective in reducing high frequency vibrations [Hernried and Lei, 1993; Inaudi

- -&-dKeuy-P9931-I- - - - - -. - - - . - -- - - - - - - - - -

I . n any case, a proper cholce of the isolation system is decisive for an adequate protection of objects and equipments [Kelly, 1999). In this paper the protection levels for the content provided by base-isolated buildings are evaluated and compared also to the protection levels provided by fixed-base structures. Three types of isolation systems &e considered, namely: (1) rubber-based isolation sys- tems, (2) steel-based isolation systems and (3) SMA (Shape Memory Alloy)-based isolation systems. The present study is based on the experimental results of a big program of shaking table tests, on reduced-scale structural models, in which the DiSGG of the University of Basilicata was involved in 1998, within the MANSIDE (Memory Alloys for Hew Seismic bolation Wvices) project [MANSIDE, 19991. Seven identical 3.3-scale, 3-storey, R/C frames were tested, including two fixed- base models and four base-isolated models, equipped with the above mentioned isolation systems. A comprehensive description of the entire program of shaking table tests a d associated results can be found ia (Dolce et al., 20011. In this pa- per, the floor response spectra obtained from the experimental storey accelerations recorded during the shaking table tests are compared, to assess the effectiveness of different isolation systems in protecting internal contents. Three Merent PGA levels are considered, approximately equal to O.l5g, 0.39 and 0.59, respectively. The internal equipment or objects, &om now on called secondary systems, are regarded as elastic single degree of freedom (SDOF) systems, having 2% equivalent viscous damping. The interaction between primary (building) and secondary system is ne-

glected;-assuinirig-a-largcdue-of the-ratio-betw&%Eiith~el~t masses. T o X i d E t 7 and better understand the experimental records, an accurate numerical model has been implemented, to simulate all the shaking table tests and compare experimental and numerical response spectra.

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Seismic Protection of Light Secondary Systems 225

2. Experimental Set up

2.1. Structural model

The tested structural models are referred to a 3-storey, Zbay, moment-resisting R/C Erame of a typical building (3.5 m interstorey height and 5 m span length in the prototype scale), as shown in Fig. l(a). Beams and columns have the same cross section (35 cm by 50 cm in the reference prototype scale) all over the structure.

The model was anchored to the earthquake platform through three heavy duty load cells, placed below each column, to measure base shear and overturning mo- ment. The isolation devices were put between the load cells and the shaking table.

The shakmg table was driven in the longitudinal direction of the hame. To avoid any out-of-plane movements of the model, two stiff steel-braced frames (one for each side of the model) were bolted to the shaking table and endowed with two pairs of low-friction sliding guides per floor.

All the structural models were defined with reference to the same full-scale prototype, so as to easily compare the results relevant to different seismic protection strategies and systems. The full-scale prototype was designed according to Eurocode 8 [CEN,1998]. Soil type B response spectrum, 0.159 PGA, "low" ductility class and 2.5 behaviour factor were selected as main design parameters. The prototype was then scaled down by a 3.3 factor, in order to fully exploit the size (4 rn x 4 m) and the payload capacity (about 150 KN) of the shaking table of the Technical University of Athens, where the tests had to be carried out.

Fig. 1. (a) k3.3 scale structural model tested on shaking table and (b) scaled acceleration protile and response spectrum of the input motion.

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226 M. Dolce & D. Cardone

Mass-similitude scaling resulted in about 77.4 kN steel blocks to be added to the model and anchored to floor slabs, to get 114.3 kN total weight.

Twenty sensors measured displacements and accelerations. Seven Endevco ac- celerometers, with f 2g range, were used to record horizontal and vertical floor wcelerations, while one accelerometer was put on the table, to record accurately the acceleration time history at the base of the model.

2.2. Seismic motion

An artificially generated acceleration profile, compatible with the response spectrum provided by EC8 for soil type B [CEN, 19981, was chosen as the test earthquake for the reference prototype. The total accelerogram duration was set equal to 20 s with 13 s stationary duration , after 2 s ramp. The acceleration profile was then scaled down in time by a factor equal to (3.3)'12, for scale consistency with the models. Figure l(b) shows the scaled ncceleration-time history and the associated 5% damping response spectrum. During the tests, the peak table acceleration was progressively increased, up to reaching the structural collapse (for the fixed-base models), or attaining the operative limits of the table or of the devices (for the base-isolated models). In order to evaluate the damage suffered by the structure, a low intensity white noise input signal (0.089) was applied to the ti~ble between one seismic test and the next.

2.3. Base isolation systems

The rubber-based isolation systems was designed according to widely used criteria, b,y assuming 1 s as target period of the base-isolated model, equivalent to 1.81 s of the prototype structure, and 10% equivalent viscous damping. Several high damp- ing rubber bearings, whose geometrical characteristics are shown in Fig. 2(a) and rubber shear modulus was approximately 0.6 N/mm2, were manufactured by TIS

Rubber-based isolation system Rubber

Fig. 2. Geometric characteristics of each rubber isolator and associated experimental force- displacement behaviour.

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Seismic Protection of Light Secondary S y s t e m 227

S.p.A. Figure 2(b) shows the mechanical behaviour of the whole rubber-based iso- lation system made of three isolators, as drawn from the experimental tests carried out at the Laboratory of the University of Basilicata. The tests were performed by applying cycles of increasing amplitude, from 5% up to 200% shear strain. at 0.3 Hz frequency, with 4 kN compression force, i.e. 1/3 ,the weight of the model, on each bearing.

In the SMA-based and steel-based isolation systems, the gravity loads are car- ried by three lubricated steel-Teflon sliding bearings. Their friction coefficient was experimentally evaluated to be of the order of 3%, thus leading to a total fric- tion force of about 3.5 kN. Additional energy dissipating capability is provided by a U-shaped steel element or, alternatively, by a SMA-based device, which also provides a strong recentring capability.

Figure ?, shows the SMA-based isolation system. It includes a re-centring iso- lation device, exploiting the superelastic properties of Nickel-Titanium SMA wires [Duerig et al., 19901, to provide up to 30 kN force and reach 100 mm displacement, and three steel-Teflon sliding bearings. The SMA wires of the device are arranged in two groups, the re-centring and the dissipating group, in order to provide the isolation system with both re-centring and additional energy dissipation capability [Dolce et al., 20001.

In the present paper two configurations of the SMA device are considered, named SMAl and SMA2 respectively. SMAl includes both the re-centring and the en- ergy dissipating group, while SMA2 includes the re-centring group only, with the same number of SMA-wire loops as SMA1. Standard design criteria lacking, the SMA-based isolation systems were simply designed to guarantee a full recentring behaviour of the entire isolation system, while transferring as low as possible force to the superstructure. Thus, the number and pre-tensioning Ievel of the ShIA wires were calibrated to provide the device with a t least 3.5 W supplemental recentring force and get double flag-shaped hysteresis loops for the complete isolation system. Accurate nonlinear analyses were then carried out, in order to check that the base displacement was compatible with the technological limits of the device, having 100 mm maximum stroke. The U-shaped steel energy dissipating element of the steel-based isolation system was calibrated to get a total force of the isolation sys- tem of the same order as that of the SMA2 system, under the design earthquake, i.e. about 10% of the model weight.

Figure 3 shows the experimental force-displacement relationships of the three configurations of SMA- and steel-based isolation systems employed during the shak- ing table tests, as obtained from the cyclic tests on the isolation devices carried out at the Laboratory of the University of Basilicata. STEEL indicates the isolation system equipped with just one U-shaped steel plate, stressed in roller bending [Dolce et al., 20001. The strongly nodinear behavioui of the SMA- and steel-based isolation systems is apparent.

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SMA and steel-based isolation systems

Sleekteffon sliding bearings

SMA2 SMA1 Steel 50 10 10

40 .a 40

30 10 30

20 10 10

I 0 10 10

5 1 " 5 o . I0 -10 .I0

.lo -10 .la

-30 -34 .10

4a 4 4

-10 -lo .10

. I n .W 40 .30 0 30 60 W 120 -120 .W .W -30 0 30 so W I10 ..la .- 40 .so 0

SMA 2: isolation system made of three steel-teflon sliding bearings and a SMA isolation with 12 "re-centring" 1 mm diameter NiTi wire loops pre-strained at about 3%. SMA 1: isolation system made of three steel-teflon sliding bearings and an isolation device based on 12 "re-centring" and 8 "dissipating" 1 mm diameter SMA wire loops pre-strained at about 3%. Steel: isolation system made of three steel-teflon sliding bearings and one U-shaped stell plate.

Fig. 3. Arrangement of the SMAIsteel-based isolation system and associated experimental force displacement behaviours.

--- - - -. - --

Figure 4 compares the secant stiffness and the equivalent damping of the four types of isolation systems under consideration. As can be seen, the SMA- and the steel-based isolation systems exhibit a very high stiffness at small displacements, which, however, rapidly decreases while increasing cyclic amplitude. As a matter of fact, at relatively large displacements, the secant s t ihess of the SMA-based

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Seismic Pmtection of Light Secondary Systems 229

-&- Rubber

* SMAl

Fig. 4. Comparison between the isolation systems in terms of (a) secant stiffness and (b) equiv- alent viscous damping, at different displacement amplitudes.

isolation system is one order of magnitude smaller than at small amplitudes. On the contrary, the secant stifiess of the rubber-based isolation system, slightly decreases while increasing cyclic amplitude and, then, it increases at very large displacements, due to the strain hardening of rubber. The equivalent damping is quite constant (about 8-9%) for the rubber-based isolation system,' while it reduces horn about 27% to about 15% when increasing cyclic amplitude, for the SMA-based systems. For the steel-based system it reaches as high values as 37% at small cyclic displace- ments, then reducing progressively, up to 30% at relatively large displacements.

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These results conhm the better control of the force transmitted to the superstruc- ture by the SMA and the steel-based systems, with respect to the rubber system. As can be noted, the activation threshold forces of the four isolation systems are quite dserent, as well a s the forces transmitted to the superstructure for a certain base-displacement, thus resulting in different shear forces, and then accelerations, acting on the models during the shaking table tests. However, SMA1 and RUBBER develop the same force (about 20 kN) at 50 rnm displacement, which approximately is the displacement that both of them reached during the tests at about 0.3g. Con- siderably lower force level (about 12 kN) and less than 50 mm displacements are reached by SMA2 and STEEL devices, during the shaking table tests at about 0.39.

3. Numerical Model

A refined numerical model was implemented to support the experimental activity, using the-finite element program DRAIN-3DX [Prakash et al., 19941. The structure was modelled as an assemblage of non-linear elements connected at nodes. The mass was lumped at the beam nodes. Geometrical dimensions, masses, constitutive laws of materials, amount and arrangement of steel reinforcement, mechanical behaviour of the devices and seismic actions were reproduced in the model with great care. Non-linear time-history analyses were carried out, by using the acceleration profiles actually recorded on the shaking table during tests.

Figure 5(a) illustrates the finite element model of the R/C frame. The "fiber" element of DRAIN-3DX was employed to model beams and columns. The cross section of each structural member was divided into a number of concrete and steel fibers, as shown in Fig. 5(a), in order to capture the effects of yielding and strain- hardening of steel, as well as of cracking, crushing and post-crushing strength loss of concrete. Each structural member was then divided into a number of fiber ele- ments, to take into account the actual arrangement of reinforcement along the R/C members. At the end of each structural element, rigid end zones were defined to simulate the actual stiffness of the beam-column joint.

In Figs. 5(b) and (c) there are reported the constitutive laws of concrete and steel. Steel bars were supposed to have the same behaviour under tension and com- pression. Both constitutive laws were constructed to simulate the mechanical be- haviour of the materials as obtained from experimental tests carried out during the construction of the structural models. The constitutive law of concrete, therefore, was speciahed to each structural model, according to the experimental outcomes.

A damping stifhess-proportional matrix, i.e. C = OK, was assumed for the -R/C-structural-members.-The-~-~oefficient-was-assumed-as-an-analysis-parameter -

and then calibrated by fitting the numerical floor displacement-time histories to the experimental ones, according to the least square method.

The SMA- and steel-based isolation systems were modelled by combining sev- eral truss and link elements (Types 1 and 9 elements) of DRAIN-3DX, having elasto-plastic or non-linear elastic behaviour. Friction in sliding bearings was taken

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Seismic Protection of Light Secondary S y s t e m 231

Fig. 5. Modelling of the R/C frame using fiber elements: (a) discretisation and constitutive laws of (b) concrete and (c) steel.

into account by a truss with a rigid-plastic behaviour. The sliding force ws taken equal to 3.5 kN, in accordance with the experimental outcomes. No viscous damp inn was m i g a d to the Gnite elements describing t.he mechanical behaviour of the -

SMA and steel isolation system. Two ditferent models were considered for rubber isolators: visco-elastic and hys-

teretic. For the visceelastic model a damping stifbess-proportiond coefficient was assumed. Stiffness and damping values were varied according to the diplacement amplitude reached. The hysteretic model was obtained by combining three elastc- plastic strain-hardening elements, assuming zero viscous damping.

In all the cases, the parameters of the numeric& models were calibrated on the experimental tests of the devices [Dolce et al., 20001.

Figure 6 compares the experimental and numerical respome of the model equipped with the four isolation systems under consideration, at about 0.39 PGA.

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M. Doke & D. Cadone

Rubber @ 031g Rubber @ 031g

Steel @ O3lg 100 - 75 - - - - - .- -- - Numcriul 50 -

Steel @ 031g

Fig. 6. Displacement-time histories and hysteresis loops exhibited by the isolation systems during test at 0.3g PGA.

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Seismic Protection of Light Secondary System 233

The comparison is made in terms of both displacement-time histories and force- displacement behaviours of the isolation systems and demonstrates the great accu- racy of the numerical model in reproducing the structural response. The results of the rubber system in Fig. 6 are referred to the visc+elastic model.

4. Results

4.1. Stmctuml response

Two preliminary remarks have to be made, before analysing the results. The first is that differences in the mechanical properties of the concrete of the structural models exist. For instance, the Young's modulus, referred to 1/3 the concrete ultimate strength, resulted equal to 22 500 Mpa for the fixed-base model, 28000 Mpa for the model with rubber bearings, and 26500 Mpa for the model with SR/IA/steel- based isolation system. The second observation is that, during the calibration of the shaking table preceding the tests on the model with rubber bearings, PGA as high as lg was accidentally generated. This caused the failure of the isolators and the damage of the R/C frame. After substituting the rubber isolators, the test program proceeded as scheduled, but on a somewhat cracked R/C frame, characterised by a natural frequency 16% less than expected.

In Fig. 7 there are reported the peak values of the recorded floor accelerations of the five models under consideration (hed-base, rubber, SMA1, SMAP and steel base-isolated), for all the PGA values experienced during the shaking table tests.

In the fked-base configuration (Fig. 7(a)), the ground acceleration is amplified at the top floor by a factor 2 to 3, in accordance with the EC8 type B elastic response spectrum, being 0.45 s the fundamental period of the prototype structure. For the model with rubber isolators (Fig. 7(b)), the peak top floor accelerations are very close to the peak table acceleration, for any seismic intensity. The same occurs in the case of steel-based isolation system at 0.319. On the contrary, in the case of SMA-based isolation system (see Fig. 7(c)) the peak top floor acceleration results of the same order of magnitude as the fixed-base model for low seismic intensities (e.g. 0.149), white it is of the same order of magnitude as the model with rubber bearings for high seismic intensities (e.g. 0.59). This has to be ascribed to the strong nonlinear behaviour of the SMA-based isolation device. It is very stiff at small displacement amplitudes (so that the structure tends to behave like a &xed-base rather than a base-isolated one), while becoming more and more flexible at Iarge displacement amplitudes, with a limited force-transmitting capacity. Obviously this behaviour can be suitably calibrated with respect to the seismic intensity level, by changing the elastic threshold force, i.e. the number as well as the prestress level of the SMA wires.

Figure 8 shows the 3rd floor acceleration time histories and the associated Fast Fourier Tkansforms (FFT) [Cooley et al., 19691 relevant to tests at about 0.3g PGA, on fixed-base and bas&solated models. As can be seen, the structural response of

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Mar acccl

PGA - 0.07g

--+--. 0.148

-..*-- 0.19g - 028g

*.-+.-.- 0.488

Mar acccl

PGA - SMA2@0.3g - Slce@0.3lg

-a- SMAl@O.l4g

--* SMAI@OJIg - SMAl@O>Og

- -- - - - - -

FB: Fixed-base model; BI-rubber: Model with rubber bearings; BI-SMA1: Model with isolation device based on re-centring and dissipating SPvIA wires; BI-SMA2: Model with isolation device based on re-centring SMA wires only; BI-steel: Model with isolation device based on steel components. Fig. 7. Peak floor accelerations for fixed-bmde and base isolated models for different PGA.

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Seismic Protection oJ Light Secondary Systems

FB: Fixed-base model; BI-rubber: Model with rubber bearings; BI-SMA1: Model equipped with isolation device based on re-centring and dissipating SMA wires; BI-SM.42: Model equipped with isolation device based on re-centring SMA wires only; BI-steel: Model equipped with isolation device based on steel components.

Fig. 8. Third floor acceleration time histories and associated FFT relevant to tests at about 0.39 PGA.

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both the fixed-base model and the model with rubber isolators is dominated by the 1st mode, though some minor peaks also occur around the second mode frequencies. For SMA and steel-based isolation systems, instead of a high and well defined peak, two squat bell-shaped curves appear. Their maximum values, occurring around the 1st and 2nd natural frequency of the linearised base-isolated structure, are very close each other and 2-3 times smalter than for the fixed-base and the rubber base- isolated structures. The strong nonlinear behaviour of the SMA and steel-based isolation systems causes the seismic energy associated to the 1st mode of vibration to concentrate not around a single frequency value but to spread over a certain frequency range, as a consequence of the continuous changes in stiffness of the isolation device (see Fig. 4).

- As observed&[K~y119&1] ,_thishishas_toPbe_eascribed t o the shape-of-the-hysteresis---- loops exhibited by the SblA and the steel-based isolation systems, and particularly to friction,-Indeed, as the motion direction changes, a sudden reversal of sign of the friction resistance in the steel-teflon sliding bearings takes place, which produces a force discontinuity inthe hysteresis Loops, as shown. in Fig. 3(c). This discontinuity generates high-frequency vibrations, that are transmitted into the isolated frame, thus exciting its 2nd mode. For the SMA devices, something like this takes place also whenever the displacement inside the device passes through zero, where an abrupt change in stiffness occurs. All that is clearly visible in the acceleration time histories shown in Fig. 8. The acceleration prose relevant to the model with rubber isolators has a sinusoidal trend, which is very similar to that of the base displacement. For the models with SMA- and steel-based isolation systems this is not true anymore, due to the superposition of a series of high-frequency waves. Also the d8erences between SMA and steel isolation systems should be noted, though they should be mainly ascribed to the lower elastic threshold force chosen for the latter system (see Fig. 3).

Obviously, the lower the seismic intensity the more conspicuous the phenomena previously described, since motion reversals become more frequent and numerous. When increasing PGA, the aspect of the FFT's relevant to the model with SMA- based isolation system (not shown in Fig. 8) tends to become similar to that of the corresponding FFT relevant to the model with rubber bearings. The second mode, indeed, becomes less and less important, while the peak associated to the first mode becomes better defined and tends to move towards lower frequency values, as a result of the increase in the amplitude of vibrations.

-- 4.2. Effects-on light -secondary systems - --

In Figs. 9-12 there are reported the 2% damping response spectra obtained from the 3rd floor acceleration time histories, for three different seismic intensities (PGA), equal to about 0.159, 0.39 and 0.59. In the diagrams, the periods have been referred to the full-scale structure, by multiplying their values by the square root of the model scale. For the fked-base model, the response spectrum relevant to 0.59 is

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Seismic Pmtection of Light Secondary S y s t e m 237

Fig. 9. 2% equivalent viscous damping response spectra relevafit to iixed-base model, as obtained kom both experimental and numerical 3rd floor acceleration time histories rewrdd during tests at different PGA's.

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BI-rubber ( 6 =2% , PGA 0.13g)

- Experimental I - HYST Numerical

BI-smal ( 5 =2%, PGA 0.16g)

0 1 2 3 4 . . - - - . - - - - - - - - - - -

T (sec)

Fig. 10. 2% equivalent viscous damping response spectra relevant to the base-isolated models, as obtained from both experimental and numerical 3rd floor apceleration time histories recorded during tests at about 0.159 PGA.

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Seismic Protection o j Lzght Secondary S y s t e m 239

81-smal (<=?a. PGA 0 . 3 1 , ~ ~

0 1 1 4

T (-1

BI-steel ( < -2%. PGA 0.33)

Enpcrimsnwl; - Nvmniul

1

Fig. 11. 2% equivalent viscous damping response spectra relevant to the base-isolated models, as obtained from both experimental and numerical 3rd floor acceleration time histories recorded during tests at about 0.39 PGA.

missing, due to the low accuracy of the acceleration-time history caused by the model collapse.

Numerical response spectra are reported in every diagram along with the corresponding experimental one. For all the models, the accordance between ex- perimental and numerical results is excellent, thus confirming the great accuracy of the numerical model in reproducing the experimental response of the whole structural system, including both the R/C frame and the isolation devices. The numerical model implemented and calibrated herein can represent a reliable tool for an extensive parametric analysis, to investigate the problem of the protection of internal contents of R/C structures.

In the response spectra relevant to the model with rubber bearings, two different numerical curves are reported. The fist (black thick line curve) has been obtained with the visco-elastic model of the isolators, the second (grey thick line curve) with the hysteretic model. The importance of choosing a proper model, to match the ex- perimental behaviour of the structure, is apparent. The visco-elastic model, which is generally preferred to the hysteretic model for Its simplicity and reliability, is not able to correctly account for the effects related to the 2nd mode of vibration of the structural system. In Table 1 there are listed the frequencies corresponding to the first three modes of vibration of the structural models under consideration,

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( < =2% , PGA 0.52g)

-- HYST Numerical '

BI-smal ( 5 =2 % , PGA OSg)

2 - 1 - Experimental /

I - Numerical I

Fig. 12. 2% equivalent viscous damping response spectra relevant to the base-isolated models, as obtained from both experimental and numerical 3rd floor acceleration time histories recorded during tests a t about 0.159 PGA.

as experimentally observed during seismic tests of increasing intensity. The values reported in Table 1 have been referred to the full-scale structure, for consistency with the response spectra previously presented. As can be seen, the natural frequen- cies of vibration of the models match the values corresponding to the peaks of the

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Seismic Protection of Light Secondary Systems 241

Table 1. Natural frequencies of vibration For each model tested.

PGA f l (Hz) f2 (Hz) f3 (Hz)

floor response spectra reported in Figs. 9-12. This clearly highlights the dangerous effects of possible tuning between equipment and structure.

For the fixed-base structure, the tuning effect takes place when the natural frequency of the content is close to the 1st natural frequency of the structure. At 0.149, this occurs at around 2.8 Hz in the model scale (1.5 Hz in full scale) and the peak acceleration experienced by the secondary system results as high as 59.

For the model equipped with rubber isolators, two apparent tuning effects occur, respectively, when the natural frequency of the content is close to the 1st and to the 2nd natural frequency of the structure. The first tuning phenomenon is characterised by a double peak, to be ascribed to the nonlinear relationship between stiffness and shear strain in the rubber isolators (see Fig. 4). As can be observed, the peak acceleration experienced by equipments with high periods of vibration, comparable with that associated to the 1st mode of the base-isolated structure (1-1.4 s in the model scale, 2-2.5 s in full scale), passes fiom lg to 2.59 while increasing PGA from 0.139 to 0.529. At the same time, equipments with low periods of vibration, comparable with that associated to the 2nd mode of the base-isolated structure (0.2 s in the model scale, 0.35 s in full scale), can record peak accelerations as great as 29.

For the model equipped with SMA- and steel-based isolation systems, only one well-defined tuning phenomenon occurs, when the period of vibration of the equip ment is close to that of the 2nd/3rd mode of vibration of the base-isolated structure (0.06-0.13 s in the model scale, 0.1-0.23 s in full-scale). In this case, the equipments experience peak accelerations of the order of 2-2-59, without significant variations while increasing PGA. It should be noted that such values are of the same order of magnitude as those relevant to the model with rubber bearings, the only differ- ence being in the width of the associated frequency range, which is larger for the SMA- and steel-based isolation systems. In the low-frequency region of the response spectra, no appreciable tuning effect is visible. The strongly nonlinear behaviour of the SMA- and steel-based isolation systems, indeed, produces a frequency shift

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242 M. Dolce €9 D. Cardone

which depends on the response magnitude. Hence, during an earthquake, the ef- fective hequency of the structural system is continuously changing, so that any equipment tuned to a specific frequency will be de-tuned during some portion of the seismic attack. As a result, the peak acceleration experienced by equipments

BI-rubber ( 5 =2% , PGA = 0.3g)

BI-srnal ( 5 =2 9% , PGA = 0.3g)

* 2 0 -

- 1 st storey /

U % vl

15 -

10 - -

- - 3rd 2nd storey /

2nd storey , I

--- 3rd-storey 4-

Fig. 13. Normalised experimental response spectra at the three floors of the models with rubber bearings (BI-rubber) and SMA-based isolation system (BI-SMAl), for PGA equal to about 0.39.

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Seismic Protection of Light Secondaq Systems 243

placed on the structure isolat,ed with ShlA or s t ee l -b~ed systems remains qGfe

conatant in a 'ide h v q range, when compared to the structure with rubber bearings. ~c~ leads to differences (up to 400%) in the low f i eWenc~ range, under low-to-moderate earthquakes-

BI-smal

_ . _ _ _ _ _ _ . . - . - -

0 r - 1

I

3 4 1 2 0 T (set)

Fig. 14. ~ ~ ~ ~ ~ h ~ ~ d exprimentd r e p o m spectra of the model with rubber bemine (BI-rubber) and S M A - M ~ ~ ~ isdrtion system (81-SMAl). while changing PG*.

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PGA = 0.159 PGA = 0.15g

PGA = 05g

J--BI-nrbberl PGA = OSg

20,

Fig. 15. Comparison between the normalised response spectra .relevant to the isolation systems under consideration, at different PGA's. The horizontal broken line corresponds to S,/PGA equal to 1.

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Seismic Protection of Light Secondary Systems 245

Figure 13 shows the experimental response spectra normalised with respect to PGA at the three stories of the structure with the rubber- and the SMA-based isolation systems, for about 0.39 PGA. As can be seen, no appreciable differences can be found at the three stories in the low-frequency region of the spectra, while the contrary holds for the high-frequency region. For both the models, the most dangerous floor turns out to be the third one.

Figure 14 shows the experimental response spectra normalised with respect to PGA, relevant to the structure with rubber bearings and with the SMA-based isolation system, for different values of PGA. As can be seen, when increasing PGA, the effectiveness of the SMA-based isolation system increases significantly and its response spectrum becomes more and more similar in shape to that associated to the rubber-based system.

Figure 15 compares the 2% viscous damping response spectra, normalised with respect to PGA, relevant to all the structural models under consideration. and for the three considered PGA's. The magnification factor S,/PGA is shown as a function of the period, on the left side, and of the frequency, on the right side, of the secondary system. The horizontal broken line drawn in each diagram corresponds to a magnification factor SJPGA equal to 1. In Table 2 there are reported the maximum accelerations experienced by four particular secondary systems (SDOF No. 1, 2, 3 and 4), each one tuned to a specific natural frequency of the primary system. In particular, the secondary systems No. 1 and 2 are tuned to the 1st naturai frequency of the fixed-base and of the rubber base-isolated models, 1.55 Hz and 0.52 Hz, respectively. The secondary systems No. 3 and No. 4, instead, are tuned to the 2nd natural frequency of the models with SMA-based and rubber-based isolation systems, 4.59 Hz and 2.76 Hz, respectively. In addition, in Table 2 there are reported the average values of the acceleration amplification factor experienced by three groups of secondary systems, having high (3-10 Hz) medium (0.8-3 Hz) and low (0.3-0.8 Hz) natural frequency of vibration.

By analysing the pad of Table 2 relevant to the four SDOF systems, the following observations can be made:

(1) At about 0.159 PGA, SDOF No. I, mounted on top of the fixed-base model, records a maximum acceleration 34 times greatcr than PGA. This value clearly expresses the risk the equipments are exposed to, if a tuning phenomenon occurs, which is often inevitable, due to the great number and variety of equipments placed in the structures of interest. With rubber-based isolation, the maximum acceleration experienced by SDOF No. 1 is around 1.5 times PGA, not much dependent on the seismic intensity in the investigated range. With SMA- and steel-based isolation systems, instead, things improve while increasing PGA. At 0.39, the maximum acceleration experienced by SDOF No. 1 is just 10% greater than PGA, in the case of steel-based isolation system, and about 20% greater than PGA, for the SMA2 isolation system. At 0.5g, then, the amplification factor reduces to 1.07 for the SMAl isolation system.

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If the scope is that of protecting equipments with frequency of vibration close to the natural frequency of the structure, at medium and high level of the seismic intensity, SMA and steel-based isolation systems result to be more ef- fective than rubber bearings. Indeed, the maximum acceleration experienced by the equipments never overcome 0.7g with the SMA isolation system, while it reached 0.889, at 0.59, with the rubber bearings. The quasi-linear behaviour of the rubber isolators leads a great amount of the input energy to concentrate in a quite narrow kequency range, where the natural frequency, of the secondary system can fall, thus giving rise to dan- gerous tuning effects. This is not the case of SMA- and steel-based isolation systems, whose strongly nonlinear behaviour implies that the input seismic en-

- . ergy ---- spreads over a wide frequency range. At 0.39, for instance, the maximum --

acceleration recorded by SDOF No. 2, placed on top of the model with rubber bearings, is about 2.5 times greater than for the SMA-based isolation systems and about 3.5 times greater than for the steel-basedkolation system. Although the problem of nonlinear secondary systems is not addressed in this paper, it is worth to underline that this concentration of energy a t low frequencies can be dangerous for the overturning of large-size heavy objects (e.g. statues, tanks, etc.), whose behaviour cannot be described by a linear SDOF. In these cases, the use of rclbber base-isolation can result even detrimental in protecting the structural contents, with respect to the fixed-base structure. The maximum acceleration experienced by SDOF No. 3 with a SMA- or steel- based isolation system is 5-8 times greater than PGA. Correspondingly, the maximum acceleration experienced by SDOF No. 4 with a rubber system is about 4 times the PGA. This emphasises the sensitivity of secondary systems tuned to the second frequency of the isolated structural systems, which re- sults to be greater for SMA- and Steetbased isolation systems, especially at low-medium intensity. However, it is worthy noting that the high-frequency vibrations transmitted by the isolation system to the superstructure do not penalise secondary systems with high fundamental frequency more than what occurs for the fixed-base structure.

The last three columns of Table 2 summarise the effects of seismic isolation on secondary systems classified in three frequency categories. The rubber-based isc- lation system appears to be the best solution for the protection of internal equip ments with high frequency of vibration (3-10 Hz). Ln presence of rubber bearings, indeed, the average amplification factor ranges horn about 1.3 to about 1.6 only.

- - Things remain fairly good f6r tKe sGelX&d isGlati6n KyiitFm,-whilG th3y get de- -

-

cidedly worse for the SMA-based isolation system. As a matter of fact, the average ampMcation factors recorded by high-frequency secondary systems, mounted on top of the model equipped with SMA-bsed isolation system, turn out to be of the same order of magnitude as those observed for the fixed-base model for the lower intensity. For secondary systems with medium frequency of vibration (0.8-3 Hz),

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the best solution seems to be the steel-based isolation system, although data are available only at 0.39 PGA. Actually, the average amplification factor is around 1.5 for the steel-based isolation system, whle it results of the order of 2 for the rubber system, regardless the PGA value. For the SMA-based isolation system, things im- prove while increasing PGA, as the average amplification factor reduces from 6.1 to 1.84, when PGA passes from about 0.159 to 0.5g. It should be observed, however, that the protection level of the equipments guaranteed by the SMA-based isolation system remains much greater than that offered by the fixed-base model, for any PGA value. For secondary systems with low natural fiequency of vibration (0.3- 0.8 Hz), the best solution turns out to be again the steel-based isolation system, as the average 1.3 amplification factor confirms. With the SMA-based isolation sys- tem, things seem to go in the same way as for the fixed-based model. In this cGe, - -- - - the average amplification factor increases while increasing PGA, passing from 1.74 at about 0.159 to 2.69 at 0.59. With the rubber bearing system, on the contrary, the average amplification factor remains quite constant while changing PGA, being of the order of 3-3.5, . i.e. - - greater than for the fixed-base model.

5. Conclusion

Although many of the conclusions of this work could already be deduced from theoretical considerations, their experimental verification was important for more consistent applications of seismic isolation, where the protection of internal contents is a primary concern.

In the experimental tests presented, the design criteria of the isolation systems were not specifically aimed at the protection of secondary systems. Nevertheless, their results have in general confirmed the effectiveness of seismic isolation in re- ducing the accelerations on the internal contents of structures. As a matter of fact the amplification factor reduced, in the average, by 2-3 times with respect to the fixed-base structure: However, the experimental results have also pointed out that each type of isolation system can result more or less effective in certain fiequency ranges, depending on their d y n d c behaviour.

Strongly nonlinear isolation systems (such a s those based on SMA or steel com- ponents), though better limiting base displacements and forces transmitted to the structure, generate high-frequency vibrations, due to the sudden changes in stiffness occurring during an earthquake. As a consequence, secondary systems with natural frequencies close to the 2nd and 3rd mode frequency of the isolated structure can be subjected to accelerations of the same order as in the fixed-base structure. On the-other-hand, quasiclinear isolation devices-(such asrubber-isolators) tend-to conr -

centrate a pea t amount of seismic energy in a quite narrow low-frequency range, in which the natural frequency of some secondary systepls can fall, thus producing undesirable tuning effects. In this case, their response may result many ( 3 4 ) times geatei'than if it were mounted on a fixed-base structure.

Generally speaking, the choice of the type of isolatiorl system, as well as its design, must be optimised with respect to the type and the dynamic characteristics

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Seismic Protection uf Light Secondary S y s t e m 249

of the internal content to be protected, as no practical design solution exists that protect any type of content in any frequency range. The following parameters should be taken into account:

r Natural frequency of the most important secondary systems to protect. 0 Initial stiffness and threshold force of the nonlinear isolation system or, alterna-

tively, natural frequency of the quasi-elastic (rubber-based) isolation system. Higher mode frequencies of the isolated structural system. Intensity of the earthquake for which the secondary system shall be preserved.

Particularly the isolation period of quasi-elastic systems and the threshold force level of strongly non-linear systems, affect significantly the response of the struc- ture, and then of secondary systems. Compatibly with other design requirements, such as the maximum admissible displacement, this undoubtedly permits to de- tune secondary systems, if their dynamic characteristics are known, and reduce significantly their response.

The numerical simulations of the experimental tests have shown the accurate re- producibility of the structural response through refined nonlinear numerical models. Their use can be reliably addressed to a specific case, when important and delicate applications of seismic isolation For internal content protection shall be made, or, alternatively, to get general design rules through extensive parametric investigation.

Acknowledgements

The experimental tests whose elaborations have lead to the results presented in this paper have been carried out within the MANSIDE BRite-Eubm Project BE95- 2168, supported by the European Commission under contract No. BRPR-CT95- 0031.

References

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Dolce, M., Cardone, D. and Ponzo, F. [2001] "Comparison of different passive control systems for R/C frames through shaking table tests," Proc. of the 5th World Congress on Joints, Bearings and Seismic Systems for Concrete Structures, Rome, Italy.

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