DEMT experimental and analytical studies on seismic isolation

Nuclear Engineering and Design 127 (1991) 409-418 North-Holland

409

DEMT experimental and analytical studies on seismic isolation F. G a n t e n b e i n a n d P. B u l a n d Commissariat g~l'Energie A tomique, Centre d'Etudes NuclOaires de Saclay, F-91191 Gif-sur-Yoette Cedex, France

This paper reviews the experimental and theoretical studies performed at C E A / D E M T related to the overall behavior of isolated structures. The experimental work consists of the seismic shaking-table tests of a concrete cylinder isolated by neoprene sliding pads, and the vibrational tests on the reaction mass of the TAMARIS seismic facility. The analytical work consists of the development of procedures for dynamic calculation methods: for soil-structure interaction where pads are placed between an upper raft and pedestals, for time-history calculations where sliding plates are used, and for fluid-structure interaction where coupled fluid and structure motions and sloshing modes are important. Finally, this paper comments on the consequences of seismic isolation for the analysis of fast-breeder reactor (FBR) vessels. The modes can no longer be considered independent (SRSS Method leads to important errors), and the sloshing increases.

1. I n t r o d u c t i o n

2.1. Structure on sliding pads

Work on seismic isolation has been performed in France for many years, and the isolation device developed by S P I E - B A T I G N O L L E S in collaboration with Electricit6 de France ( E D F ) has been incorporated in the design of pressurized-water reactor (PWR) nuclear power plants [1-3]. Related to the development of isolation devices, experimental and analytical studies have been performed in C E A / D E M T and will be reviewed in this paper. These studies were generally identified by the designer or the utility. Thus, they were not part of a general program. In addition to the review, work that needs to be performed especially for fast-breeder reactors will be briefly presented. As a preface to the detailed review of the studies, one must note that as for nonisolated nuclear power plants, the isolated nuclear power plant analysis is separated into two parts: (a) study of the " b u i l d i n g " for which it is necessary to take into account the isolation device behavior; and (b) study of the secondary system (equipment, etc.) for which it may be necessary to take into account the particular shape of the floor-response spectrum.

First of all, seismic tests on a concrete cylinder (height = 4.4 m) isolated by sliding pads were performed on the V E S U V E shaking table [4]. The cylinder (mass = 13.6 t) represents a ~ - s c a l e reactor building and is placed on four bearings (fig. 1). The natural frequency of the structure is 2.3 Hz. Horizontal seismic motions ( T A F T 1952 and EL C E N T R O 1940) are applied to the shaking table. Taking into account a scale factor (frequency shift with a coefficient equal to 2.5), the maximum acceleration of the shaking table is 0.6 g. Calculations performed by S P I E - B A T I G N O L L E S have shown good agreement with test results for the slippage and the cylinder accelerations [1,2].

2. DEMT

Experimental

studies

A few experimental studies related to the dynamic behavior of isolated structures and the validation of analytical methods were performed at C E A / D E M T . 0029-5493/91/$03.50

2.2. Structure on springs

Another isolated structure studied by D E M T is the reaction mass of the new C E A / D E M T seismic facility T A M A R I S [4] (fig. 2). The four shaking tables are grouped in a unique 2500 t reaction mass (16 x 16 × 7.25 m) which is suspended by helicoidal springs and viscous dampers to prevent vibration transmission. Tests were performed before installation of dampers to check calculations. The measured frequencies of the reaction mass are as follows: - 1.0 Hz for horizontal translation - 1.3 Hz for rotation about vertical axis - 1.4 Hz for vertical translation - 1.5 Hz for rotation about horizontal axis [5]. Experimental results for frequencies are slightly

© 1991 - E l s e v i e r S c i e n c e P u b l i s h e r s B.V. ( N o r t h - H o l l a n d )

410

F Gantenbein, P. Buland / Studies at D E M T (France)

higher than calculated. This discrepancy is related to the calculation model because, during these tests, the reaction mass was not equipped with all its components (shaking tables, etc.) while they were taken into account in the calculation.

3. D E M T

- analytical studies

The analytical studies performed are mainly related to the development of methods for the analysis of the soil-structure interaction in case of many pedestals, the

behavior of sliding structures, and the fluid-structure in teraction. 3.1. Soil-structure interaction

In the case of seismic isolation using sliding (or nonsliding) elastomer bearing pads, the foundation includes an upper raft and pedestals. The pads are placed between this upper raft and the pedestals which are connected to a lower raft. The horizontal flexibility is due mainly to the horizontal flexibility of the pads (if the soil is sufficiently

Fig. 1. Seismic test on a concrete cylinder isolated by sliding bearing pads.

c~

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F. Gantenbein, P. Buland / Studies at D E M T (France)

stiff). The influence of soil-structure interaction and of the pedestal flexibility is then very small. For the vertical motion, the influence of the pedestals and, consequently, of the soil deformation in the vicinity of the pedestals was studied. The stiffness for the lower raft was taken to be the same as that for the soil. The impedance calculation was performed by the computer code SOPHONIE [6], which uses the Deleuze formulation [7] for calculating the motion in any place on the surface of a homogeneous soil when a force is applied on a circular foundation. Such a calculation could have been performed by using the ATC methodology [8]. A hexagonal raft was considered and the pedestals were placed in a hexagonal net. The influence of the size

of the upper raft on the impedance was studied while maintaining the distance between the pedestals and the cross-sectional area of the pedestals constant. Thus, the number of pedestals is closely related to the size of the raft. The calculations show that (fig. 3): (a) vertical stiffness of the isolated raft foundation is similar to the impedance of a conventional raft foundation with the same surface area, since the number of pedestals is significant (or the surface is great); (b) damping of the isolated raft foundation is equal to 0.75 7; ~ is the damping of the conventional raft foundation, with the same surface area. In conclusion, for base-isolated nuclear power plants

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413

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where the number of pedestals is very important, vertical impedance can be estimated from the usual formula since damping is often limited (vertical damping is always very important for a raft of this size). 3.2. Calculation o f sliding structures

The usual linear method with oscillator spectrum is no more applicable for structures isolated by sliding bearing devices. In this case, nonlinear time-history calculations must be performed to determine the slippage and the response of the building (stresses-floor-response spectrum, etc.). An interesting procedure to perform such nonlinear calculations is based on the modal analysis technique [9,10]: (a) a modal description of the free-free structure (without boundary conditions between the raft and the foundation) is considered; (b) the nonlinear response is calculated from the solution of uncoupled equations (modal amplitude equations) in which external forces associated with the sliding or nonsliding phases are added to the seismic forces. Simple structures have been studied by this method: (a) The floor-response spectra calculated for a sliding mass (friction coefficient = 0.1) and a nonsliding mass are presented in fig. 4. For the nonsliding

mass, this spectrum is the spectrum of the support motion; it is a filtered US Nuclear Regulatory Commission (USNRC) spectrum. The filter is equivalent to an oscillator with a frequency of 3 Hz and 5% damping. One can notice that the sliding leads to a decrease of the spectrum equal to 4 for frequencies around the peak and to 2 for higher frequencies. (b) A two-degrees-of-freedom system has also been studied (an oscillator fixed on a sliding mass) for various values of the masses and of the friction coefficient. The maximum acceleration of the masses in case of a sliding mass equal to half of the oscillator mass (r = 0.5) is presented in figs. 5 and 6 for a support motion represented by the San Francisco 1957 accelerogram. These results show that as the sliding increases, the acceleration of the oscillator mass decreases, and that the effect of sliding on the sliding mass motion is more complex. The influence of the ratio r [11] is similar to what was observed in ref. [12]: the oscillator acceleration for a given friction coefficient decreases generally as the ratio r decreases, whatever may be the oscillator frequency. Additional information on the complexity of the sliding mass motion is given from its floor-response spectrum (fig. 7). It shows that for a low friction coefficient, the spectrum is much lower than the support spectrum except around the second mode of the free

414

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system (the first mode is a body-motion mode, the second mode corresponds to out-of-phase motion of the two masses), f2 = f 0 v / O + r ) / r , which is, in this case, equal to 10.4 Hz. This nonlinear dynamic analysis method can, of course, be applied for elastomer sliding pads. In the grip phase, the pads act like springs with a noninfinite stiffness. For nonlinear structures, it is necessary to perform time-history calculations. To date, it does not seem evident to define some general equivalent linear

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model which could avoid these nonlinear time-history calculations; additional damping [13] does not allow us to reproduce the shape of the fig. 7 spectrum. 3.3. F l u i d - s t r u c t u r e interaction

The modal characteristics of structures in air are modified when they are immersed in a dense fluid; the fluid modifies the mass matrix of the structure by added masses and mass coupling. In addition, when there is a free surface, the gravity effect on this surface introduces some additional modes called sloshing modes. For FBRs, it is necessary to precisely estimate these modes as they are in the frequency range of the isolated building: lower than 1 Hz for the sloshing modes and for the first modes with fluid-structure coupling [14,15]. A general method to compute fluid-structure interaction has been developed and implemented in the computer codes of the C A S T E M system [16] and was validated by analytical and experimental results on simple [16,17] and on more complex [18] structures. It solves the fluid and the solid (shells, etc.) equations simultaneously; variables related to the fluid pressures are the additional unknowns for fluid. An example of the results obtained for a tank filled with water is given in fig. 8.

415

F. Gantenbein, P. Buland / Studies at DEMT (France) Normalized pressure of water

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4. Influence of base isolation on F B R seismic analysis

4.1. Sloshing

SuperPhenix 2 (SPX2) has been studied with isolation devices for both horizontal and vertical motions because an FBR primary coolant system is very sensitive to seismic loads [19]. During the preliminary phase of the design, it was pointed out that additional work had to be done to validate the seismic analysis methods. Of course, the effects of isolation on the design of the vessels and components will strongly depend on the choice of the first natural frequency of the isolated structure (that is to say, on the pad stiffness and on their number) and on the damping of the isolation device. Nevertheless, general comments on what should be done to improve the design methodology can be provided and the required research programs are discussed in the framework of the European agreement [20,21].

For nonisolated reactors, the sloshing modes have a small contribution because the associated spectrum contains relatively small amplitudes at low frequencies. In case of isolation, the spectrum at the base of the building is greater than the soil spectrum at low frequency, and the wave height is increased. Therefore, it seems necessary to: (a) make a calculation of the wave motion taking into account the detailed three-dimensional (3D) geometry of the fluid contained in the internal vessel (methods similar to that mentioned in the previous chapter can be applied); (b) determine by tests the damping of such modes. Studies on the influence of the mockup scale on the measured damping should be performed; (c) verify that the wave motion does not introduce

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F. Gantenbein, P. Buland / Studies at DEMT (France)

nonlinearities, for example, by impacting on walls, and eventually to develop methods to take the nonlinearities into account. 4.2. Modal analysis of the reactor block

Usually, the modal analysis of the reactor block is performed using axisymmetric models. The components are represented by equivalent axisymmetric structures.

The seismic response calculated by this two-dimensional (2D) model has been compared to 3D calculations (fig. 9) for nonisolated reactors [15]. While the modal basis calculated on the two models is different, the discrepancies between the seismic responses do not exceed 30%. In the case of isolated reactors, the contributions of the various modes in the global responses of two models are different. The lowest modes which correspond to a strong coupling between the fluid and the

Fig. 9. 3D model of FBR internals.

F. Gantenbein, P. Buland / Studies at DEMT (France)

vessels have a higher c o n t r i b u t i o n in the 3D calculations. Since the d y n a m i c pressures o n the shells are i m p o r t a n t for the design of these shells (buckling), the validity of the 2D calculations must b e verified. 4.3. Core behavior

A n iterative procedure is performed to study the seismic b e h a v i o r of the core [22]. First, a linear model of the core is i n t r o d u c e d in the reactor block model. The seismic analysis of the reactor block gives the diagrid motion. A n o n l i n e a r analysis of the core is then performed for the n o n l i n e a r b e h a v i o r of the core. This analysis defines a new equivalent linear model of the

417

t r u m m e t h o d is the value of the " t r u n c a t i o n frequency," a frequency above which the m o d a l c o n t r i b u t i o n is static so all the m o d e s are in phase. T h e second p r o b l e m is validity of the spectrum m e t h o d applied to a structure whose frequencies are out of the r e s o n a n t part of the spectrum. A modification of the C Q C m e t h o d has b e e n proposed in the case of a filtered excitation [25] w h e n the filter is an oscillator. F o r general cases where the i n p u t m o t i o n is more complex, some i m p r o v e m e n t of the spectrum m e t h o d must be studied, starting from the basic r e f i n e m e n t equations of the oscillator response to r a n d o m process [26].

core.

This procedure has been tested extensively for nonisolated reactors. It should b e validated for the isolated reactors, because the peak spectrum frequency will be lower t h a n the frequency of the fuel element, c o n t r a r y to the condition that prevails in a nonisolated reactor. 4. 4. Seismic calculation of the reactor block

Two m e t h o d s are generally used for seismic analysis of linear structures such as the reactor block of FBRs. The first is the time-history method. This can be perf o r m e d o n the finite-element model or b y modal synthesis. (The modal amplitudes are calculated for each time step.) This m e t h o d requires knowledge of the soil motion (or at least the raft motion), a n d it is r e c o m m e n d e d practice to use several time histories. However, the time-history m e t h o d is expensive. The second m e t h o d is the modal c o m b i n a t i o n m e t h o d with oscillator spectrum. This well-known m e t h o d is easy to apply a n d not expensive, so it is the most c o m m o n method. Improvem e n t s to the SRSS c o m b i n a t i o n are available [23,24] a n d are necessary w h e n the modes have closely spaced frequencies so that the m o d e coupling c a n n o t be neglected. It is necessary to recall that in application of the s p e c t r u m m e t h o d to buildings, the spectrum has a large frequency b a n d w i d t h (f~(0.5, 20 Hz)), the main frequencies of the building are in the amplified part of the spectrum, a n d the higher mode c o n t r i b u t i o n is estim a t e d from a " s t a t i c " calculation. These higher modes have a forced response to the acceleration. In the case of a n isolated FBR, the floor spectrum at the base of the raft is the i n p u t for the reactor block analysis. It has a p r e d o m i n a n t peak a r o u n d 1 Hz, a n d there are small amplifications a r o u n d 4 Hz (in the frequency range of the reactor block i m p o r t a n t modes) [25]. The first p r o b l e m to solve before using the spec-

5. Conclusion Seismic isolation offers a n effective way to reduce the seismic loads o n the p r i m a r y coolant structures. D e v e l o p m e n t a n d validation of n u m e r i c a l m e t h o d s have already been performed, b u t some work is still needed for the FBR.

References [1] F. Jolivet and H. Richli, Aseismic Foundation System for Nuclear Power Stations, Trans. 4th SMiRT Conf., San Francisco, CA, paper K9/2 (1977). [2] R. Gueraud, J.P. Noel Leroux, M. Livolant and A.P. Michalopoulos, Seismic Isolation Using Sliding-Elastomer Bearing Pads, Nucl. Eng. & Des. 84 (1985) 363-377. [3] C. Colladant, Seismic Isolation of Nuclear Power Plants: EDF's Philosophy, 1st Post SMiRT Conf. Seminar on Seismic Base Isolation of Nuclear Power Facilities, San Francisco, CA, 1989 (in this issue). [4] P. Buland, C. Berriaud and J.C. Queval, TAMARIS - A New Seismic Facility at CEA/SACLAY, Trans. 9th SMiRT Conf., Lausanne, Switzerland, paper K/19 (1987). [5] P. Buland, C. Berriaud, and P. Jamet, TAMARIS - A New European Facility for Low-Frequency Multi-Axis Vibration Simulation, Noordwijk Conf., 1988. [6] F. Gantenbein, Influence of Foundation Deformation on Seismic Response of PWR Reactor Building, Trans. 9th SMiRT Conf., Lausanne, Switzerland, paper K / 2 0 (1987). [7] G. Deleuze, Rrponse a un Mouvement Sismique d'un I~lifice Pos6 Sur un Sol I~lastique, Annales de l'Institut Technique du B~timent et des Travaux Publics, no 234 (June 1967). [8] Tentative Provisions for the Development of Seismic Regulations for Buildings, ATC-3-06 Report, Chap. 6, pp. 388-389. [9] D. Brochard, F. Gantenbein, and M. Petit, Linear and Nonlinear Applications of Modal Analysis to Seismic Loadings, PVP Conf., Honolulu, HI, 1989.

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1~ Gantenbein, P. Buland / Studies at D E M T (France)

[10] D. Brochard and F. Gantenbein, Seismic Analysis of Sliding Structures, Trans. 10th SMIRT Conf., Los Angeles, CA (1989). [11] D. Brochard and F. Gantenbein, Analyse Sismique des Structures Glissant Sur Leur Base, 2~me Colloque National de Grnie Parasismiqine, St.-Remy-les-Chevreuse, (April 1989). [12] M. Mostaghel and J. Tambakuchi, Response of Sliding Structure to Earthquake Support Motion, Earthquake Eng. and Struct. Dyn. 11, (1983) 729-748. [13] S.H. Crandall, S.S. Lee and J.H. Williams, Accumulated Slip of a Friction-Controlled Mass Excited by Earthquake Motions, ASME Trans. J. Appl. Mech. 41 (1974) 10941098. [14] P. Descleve and J. Dubois, Analyses of the Dynamic Behavior of Nuclear Power Reactor Components Containing Fluid, Trans. 5th SMIRT Conf., Berlin, paper B4/3 (1979). [15] F. Brabant, F. Gantenbein and R.J. Gibert, 3D Dynamic Fluid Structure Coupling Analysis of Pool Type LMFBR Structures by Substructuring Method, p. 128, 5th PVP. Conf., San Diego, CA, 1987. [16] F. Jeanpierre, R.J. Gibert, A. Hoffmann and H. Livolant, Description of a General Method to Compute the FluidStructure Interaction, Trans. 5th SMiRT Conf., Berlin, paper B4/1 (1979). [17] M. Dostal, P. Descleve, F. Gantenbein and L. Lazzeri, Benchmark Calculations on Fluid Coupled Coaxial Cylinders Typical to LMFBR Structures, Trans. 7th SMIRT Conf., Chicago, IL, paper B8/6 (1983). [18] A. Epstein, R.J. Gibert, F. Jeanpierre, M. Livolant and R. Assedo, Hydroelastic Model of PWR Reactor Internals, SAFRAN 1, Keswick Conf., 1978. [19] G. Feuillade and P. Richard, Evolution des Dispositifs

[20]

[21]

[22]

[23]

[24]

[25]

[26]

Parasismiques du Batiment Reacteur Projet SPX2 Depuis 1980 et Cons6quences Sur les Chargements du Bloc rracteur, ler Colloque National de Grnie Parasismiqine, St. Remy-les Chevreuse, Jan. 1986. M. Dostal, A. Martelli, A. Bernard, F. Gantenbein, K. Peters and A. Preumont, Research and Development Studies on Plant and Core Seismic Behavior for the European Common Fast Reactor, Conf. on the Science and Technology of Fast Reactor Safety, Guernsey, May 1986. Proceedings of the Specialists' Meeting on Fast-breeder Reactor-block Antiseismic Design and Verification, International Working Group on Fast Reactors (IAEA), 1WGFR-65, Feb. 1988. D. Brochard, F. Gantenbein and R.J. Gibert, Seismic Behavior of LMFBR Cores, Trans. 8th SMiRT Conf., Brussels, Belgium, paper EX2/2 (1985). B.F. Maison, C.F. Neuss and K. Kasai, The Comparative Performance of Seismic Response Spectrum Combination Rules in Building Analysis, Earthquake Eng. and Struct. Dyn. 11 (1983) 623-647. T. Fakhfakh and R.J. Gibert, Statistical Analysis of the Maximum Response of Multi-Degree System to a Seismic Excitation, Trans. 9th SMiRT Conf., Lausanne, Switzerland, paper K7 (1987). J. Betbeder-Matibet and K. Christodoulou, Calcul Sismique des Cuves Avec Interaction Fluide-Structure, 1er Colloque National de Grnie Parasismique, St.-Remy-lesChevreuse, Jan. 1986. M. Livolant, F. Gantenbein, and R.J. Gibert, Mrthodes de Calcul Basres sur l'Utilisation des Spectres Sismiques Fondements Statistiques, Grnie Parasismique - Presses des Ponts-et-Chaussres, Ch. V. 8.