Response Spectrum Method in Seismic Analysis and Design of Structures

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Response Spectrum Method in Seismic Analysis and Design of Structures

AJAYA K U M A R GUPTA. Prokssor of Civil Engineering North Carolina State University F O R E W O R D BY WILLIAM J. H

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Response Spectrum Method In Seismic Analysis and Design of Structures

AJAYA K U M A R GUPTA. Prokssor of Civil Engineering North Carolina State University


WILLIAM J. H A L L Professor and Head. Civil Engineering Universit.~of lllinois at Urbana Champaign



Q 1990 by


Blackwell Scientific Publications, Inc. Editorial offices: 3 Cambridge Center. Suite 208 Cambridge, Mas~gchusetts02142, USA Osney Mead, Oxford OX2 OEL, England 25 John Street, London WClN 2BL, England 23 Ainslie Place, Edinburgh EH3 6M, Scotland 107 Bany Street. Carlton ' Victoria 3053, Australia "



All rights reserved. No part of this book may be reproduced in any form or by any electronic or mechanical means, including information storage and retrieval systems, without permission in writing from the publisher, except by a reviewer who may quote brief passages in a review First published 1990 Set by Tima Graphics, Singapore Printed and bound at the University Press, Cambridge. England 90 91 92 93 4 3 2 1

USA Blackwell Scientific Publications, Inc. Published Business Services PO Box 447 Brookline Village Massachusetts 02 147 (Orders: Tel: 61 7 524-7678) Canada Oxford University Press 70 Wynford Drive Don Mills Ontario M3C 159 (Orders: Tel: 416 441-2941) Australia Blackwell Scientific Publications (Australia) Pty Ltd 107 Barry Street Carlton, Victoria 3053 (Orders: Tek 03 347-0300) Outside North America and Australia Marston Book S e ~ c e Ltd s PO Box 87 Oxford OX2 ODT (Orders: Tel: 01 1 44 865 79 1155) Library of Congress Cataloging-in-Publication Data Gupta Ajaya K. Response spectrum method in seismic analysis and design of structures / Ajaya Kumar Gupta; foreword by W.J. Hall. (New directions in cm. Pcivil engineering) ISBN 0-86542-1 15-3 1. Earthquake engineering. 2. Structural engineering. 3. Seismic waves. 1. Title. 11. Series. TA654,6.G87 1990 624.1'7626~20


British Library Cataloguing-in-Publication Data Gupta, Ajaya Kumar Response spectrum method in seismic analysis and design of structures, I. Structure. Analysis I. Title 11. Series 624.1'7 1 ISBN 0-632-02755-X


A .


New Dirdons

in Civil Engineering

S E R I E S E D C T Q ~W.F.

CH E N Purdue University

Dedicated to my parents Dr Chhail Bihari Lal Gupta and Mrs Taravali Gupta


Foreword, ix Preface, xi Acknowledgments, xv 1 Structural dynamics and response spectrum, 1 I. 1 Single-degree-of-freedom system, I

1.2 Response spectrum, 2 1.3 Characteristics of the earthquake response spectrum, 6 1.4 Multi-degree-of-freedom systems. 7 References, 10

2 Design spectrum, 11 2.1 Introduction, I I 2.2 'Average' elastic spectra, 12 2.3 Site-dependent spectra. 16 2.4 Design spectrum for inelastic systems, 23 2.5 Comments. 27 References, 28

3 Combination of modal responses, 30 3.1 3.2 3.3 3.4

Introduction, 30 Modes with closely spaced frequencies, 31 High frequency modes-rigid response, 39 High frequency modes-residual rigid response, 45 References. 49

4 Response to multicomponents of earthquake, 51 4.1 4.2 4.3 4.4 4.5 4.6

Introduction, 5 1 Simultaneous variation in responses, 52 Equivalent modal responses, 55 Interaction ellipsoid, 59 Approximate method, 60 Application to design problems, 62 References, 64

5 Nonciassically damped systems, 66 5.1 Introduction, 66 5.2 Analytical formulation, 67 5.3 Response spectra. 7 1


5.4 5.5 5.6 5.7 5.8

Key frequencies f L and f H , 74 Modal combination, 75 Modal combination for high frequency modes, 77 Modal combination for high frequency modes-residual Application, 8 1 References, 87

rigid response, 78

6 Response of secondary systems, 89 6.1 6.2 6.3 6.4 6.5 6.6 6.7 6.8 6.9 6.10 6.1 1

Introduction, 89 Formulation of the coupled problem, 9 1 Coupled modal properties, 95 Coupled response calculation, 98 Comparison of coupled response with the response from conventional IRS method, 101 An alternate formulation of the coupled response, 106 Secondary system equivalent oscillators. 108 Evaluation of instructure spectral quantities, 1 10 Examples of instructure response spectra, 1 I4 Correlation coefficients, 1 16 Response examples, 1 18 References. 124

7 Decoupled primary system analysis, 125 7.1 7.2 7.3 7.4

Introduction, 125 SDOF-SDOF system, I26 MDOF-MDOF systems, 130 Application of the frequency and response ratio equations, 131 References. 138

8 Seismic response of buildings, 139 8.1 8.2 8.3 8.4

Introduction, 139 Analysis, I39 Building frequency, 144 Seismic coefficient, 144 References. 152

Appendix: Numerical evaluation of response spectrum, 153 A. I A.2 A.3 A.4 A.5

Linear elastic systems, 153 Bilinear hysteretic systems, 156 Elastoplastic systems, 158 Notes for a computational algorithm. 159 Records with nonzero initial motions, 160 References, 163

Author index, 165 Subject index, 167


This book devoted to the Response Spectrum Method contains concise sections on a number of the major topics associated with the application of spectrum techniques in analysis and design. Although the theory of spectra has been understood for some extended period of time, it was only in the past twenty years that the approach was adopted in a major way by the profession for use in engineering practice. This development came about as a result of three major factors, namely that the theory and background of spectra was more fully understood, that the theory was relatively simple to understand and use, and because there was a need for such a simple approach by the building codes and by the advanced analysis techniques needed in the design of nuclear power plants and lifeline systems. The author rather directly presents his interesting and informative interpretations of various spectrum techniques in the topical chapters. He correctly points out that much work remains to be accomplished, which is accurate, for spectra in general only depict maxima of various effects, and in many cases, especially where nonlinear effects are to be treated, it is often desirable to know more about the response than just a maximum value. Research on such topics presently goes forward on such matters at a number of institutions, and in time will lead to even greater understanding of the theory, and to new approaches of application. In this connection one can cite subtle yet important differences in use and interpretation of spectra. For example, the term 'response spectrum' normally is used to refer to a plot of maximum response parameters as a function of frequency or period, for a given excitation of the base of a single-degree-of-freedomdamped oscillator, as for acceleration time history of excitation associated with a specific earthquake. On the other hand a design spectrum is a similar shaped plot selected as being representative of some set of such possible or plausible excitations for use in design; as such it is a characterization of effects that might be expected as a result of some possible range of excitation inputs, and possibly adjusted to reflect risk or uncertainty considerations, personal safety requirements, economic considerations, nonlinear effects, etc. One can immediately discern the differences, directly or subtly as may be the case. It is believed that the reader will find the interesting presentation by Dr Ajaya Gupta to be educational and informative, and hopefully such as to promote additional effort to improve even further our understanding of the theory and applications thereof. W. J. HALL Professor and Head, Civil Engineering University of Illinois at Urbana-Champaign

In modern earthquake engineering the response spectrum method has emerged as the most commonly used method of analysis. The primary reason for this popularity is the fact that it provides the designer with a rational and simple basis for specifying the earthquake loading. Another reason often cited is that the method is computationally economical. If a comparison is made between the computational effort required in, say, a modal superposition analysis of a multidegree-of-freedom structure subjected to a specified ground motion history, and that in a response spectrum analysis including the evaluation of the response spectrum from the same motion history, it is not clear whether the response spectrum method would do much better. A major part of the effort, which is common in both the methods, is the solution of the eigenvalue problem. In fact, if the objective is to evaluate the response of a structure subjected to a known earthquake ground motion, there should not be any question about using a standard time-domain analysis, or alternatively, an equivalent frequencydomain analysis. It is when we are designing a structure for a potential future earthquake that the response spectrum method is much more relevant. Criticisms of the response spectrum method arise from the fact that thetemporal information is lost in the process of evaluating the spectrum. In the words of Robert Scanlan:' 'Multi-degree-of-freedom cases are thus improperly served, intermodal phasings, in particular, being unaccounted for.' Further he points out: 'The needs arising in the design of secondary responding equipment (piping, machinery, etc., on upper floors of a structure) are not adequately met by the given design response spectra. That is, the given primary shock spectra do not lead directly and simply to definition of corresponding secondary shock spectra.' Similar difficulties arise in combining the responses from three components of the earthquake. Much progress has been made in the last decade. Lack of temporal information in the response spectrum method no longer appears to be a handicap. Rational rules are now available to combine responses from various modes, and from three components of earthquake motion. These rules account for the physics of the problem, and can be further justified in the same spirit as the design spectrum itself, as a representation of expected response values in an uncertain world. Response of secondary systems can now be evaluated using efficient modal synthesis techniques in conjunction with the response spectrum method. Alternatively, the secondary spectrum, or the instructure response 'R. H. Scanlan. On Earthquake Loadings for Structural Design, Earthquake Engineering and Sfrucfural Dynamics, Vol. 5, 1977, pp. 203-205

spectrum can be evaluated by applying similar modal synthesis techniques to the secondary single-degree-of-freedom oscillator coupled to the primary system. These new techniques directly use the design response spectrum at the base of the primary structure as seismic input and account for the effects of mass interaction (between the equipment and the structure) and of multiple support input into the secondary system. In doing so it is no longer necessary to convert the design response spectrum into a 'compatible' motion history or a power spectral density function. The question of noncIassica1 damping introduced in the coupled primary-secondary system, which had not even been specifically raised 10 or 15 years ago, is now adequately addressed. This brings us to the objective of this book. It is intended to bring together in one volume the wealth of information on the response spectrum method that has been generated in recent years. Needless to say that this information has reached a critical mass suitable for a book. This book can be used as a text or as reference material for a graduate level course. Although Chapter 1 begins with the introductory information about the single-degree-of-freedom systems that leads into the definition of the response spectrum, 1feel that most students will be more comfortable with the material in subsequent chapters if they already have had an introductory structural dynamics course. This book should also serve as a useful reference for prac~icingengineers. It should help them appreciate the analytical techniques they are already using. In many cases the book may also help them improve those techniques, especially when the improvement would lead to enhanced accuracy, often resulting in significantly lower response values. It is assumed throughout the book that we are dealing with linear systems. There are two exceptions. In Chapter 2 a brief treatment is given to inelastic response spectra. Chapter 8 deals with conventional buildings which are customarily designed to undergo significant inelastic deformation under the worst loading conditions. Inelastic behavior has always been a part of seismic design of buildings, unintentionally in the beginning, and later with full knowledge and intention. Yet, our knowledge of the topic is relatively limited. Inelastic seismic behavior and design continue to be a topic of active research. Detailed coverage of current research on the topic goes beyond the realm of the response spectrum method, and is beyond the scope of this book. Brief treatments in Chapter 2 and Chapter 8 are intended to provide a useful link between the response spectrum method and the design of conventional buildings. It should be of particular interest to the students to see the link established and, at the same time, recognize the limitations of the link. I have emphasized deterministic modeling of the earthquake response phenomenon. For a given earthquake ground motion, the maximum response values for a single-degree-of-freedom system-which are the basis of the definition of response spectrum-are deterministic quantities. For a multidegree-of-freedom system, therefore, the maximum response values in individual modes are also deterministic quantities. The modal combination rules are based

partly on the physics of the problem, that is on deterministic concepts, and partly on the random vibration modeling of the phenomenon. Strictly speaking, then, these rules do not apply to responses from individual earthquakes. On the other hand, we can look upon the modal combination rules as tools for giving approximate values of the deterministic maximum response values. It is in this spirit that the response spectrum analysis results have been repeatedly compared with the corresponding time-history maxima for individual earthquakes, treating the latter as the standard. This concept is especially powerful when judging two or more modal combination rules within the response spectrum method. A rule which models the physics well is likely to give results which are reasonably close to those obtained using the time-history analysis. Probabilistic concepts play an important role in the definition of the design spectrum, as they do in defining other kinds of loads too. These concepts are most useful when all the available deterministic tools have been carefully employed. One should not replace the other. Great strides have taken place in recent years in the development and application of random vibration techniques to the earthquake response problems. Important contributions have been made to the response spectrum method using the random vibration concepts. This book has not covered those techniques and concepts for most part. My interest in the response spectrum method has been the primary motivation for writing this book. This interest has been sustained through many years of research on related topics in collaboration with coworkers and students. Such personal involvement in the topic has its advantages and disadvantages in writing a book. The advantages are obvious. The main disadvantage is that I may not be able to do full justice in presenting the works of other researchers. To that end, I shall welcome criticism and suggestions from the readers, which I hope will improve the future editions of this book. A. K. GUPTA


My interest in the response spectrum method started during my years at Sargent and Lundy in Chicago (1971-76). My division head, Shih-Lung (Peter) Chu, asked me to work on the combination of responses from three components of an earthquake. A former graduate student colleague from the University of Illinois at Urbana-Champaign, Mahendra P. Singh (now at Virginia Polytechnic Institute and State University) was also a coworker at Sargent and Lundy and was among those who willingly shared their knowledge. During my association with Illinois Institute of Technology (1976-80), I joined the American Society of Civil Engineers (ASCE) Working Group charged with preparing a Standard for Seismic Analysis of Safety Related Nuclear Structures. Robert P. Kennedy, who chaired the effort, encouraged me to become involved in the combination of modal responses. Another colleague in the group, Asadour H. Hadjian from Bechtel, Los Angeles actively participated in the resolution of the topic. I came to North Carolina State University in 1980 and have had a series of students who have participated in the efforts related to the response spectrum method. Karola Cordero and Don-Chi Chen worked on the modal combination methods. The ASCE Working Group was deliberating on developing the criterion for decoupled analysis of primary systems (198 1) when I became interested in the topic along with another former student Jawahar M. Tembulkar. The decoupling study serendipitously led me and Jing-Wen Jaw into the coupled response of secondary systems (1983). Jerome L. Sachman and Armen Der Kiureghian were very helpful in keeping us informed about the related developments at the University of California at Berkeley. Min-Der Hwang and Tae-Yang Yoon are present graduate students who have helped in this project in many ways. Ted B. Belytschko, of Northwestern University and an editor of Nuclear Engineering and Design,has been responsible for the publication of many of our papers. He also reviewed early outlines of the present work, suggesting valuable improvements. William J. Hall of the University of Illinois; Robert H. Scanlan of the Johns Hopkins University; Bijan Mohraz of Southern Methodist University and formerly my graduate advisor at the University of Illinois (1 968-7 I); Takeru Igusa of Northwestern University; and Vernon P. Matzen, James M. Nau, Arturo E. Schultz and C. C. (David) Tung, my colleagues at North Carolina State University, have read all or part of the manuscript and offered valuable comments. It has been a pleasure to work with Blackwell Scientific Publications, in particular with Navin Sullivan, Edward Wates and Emmie Williamson. W. F.

Chen of Purdue University, Editor of the series New Directions in Civil Engineering, facilitated prompt review of the manuscript. The manuscript was produced by Engineering Publications at North Carolina State University under the direction of Martha K. Brinson, who was assisted by Sue Ellis and Kraig Spruill in word processing and by Mark Ransom and his coworkers in preparing illustrations. My talented and beautiful daughters Aparna Mini and Suvarna (Sona) gave me their unconditional love and support. To them, to everyone named above and to the many other coworkers and students who have assisted me on various occasions, I acknowledge a deep sense of gratitude.

Chapter l/Structural dynamics and response spectrum

1.1 Single-degree-of-freedom system Figure I . l(a) shows an ideal one story structure model. It has a rigid girder with lumped mass m which is supported on two massless columns with a combined lateral stiffness equal to k. The energy loss is modeled by a viscous damper, also shown in the figure. This structure has only one degree of freedom, the lateral displacement of the girder. Under the action of the earthquake ground motion, u,, the structure deforms, Figure l.l(b). The relative displacement of the girder with respect to the ground is u. The total displacement of the girder is u-(- u,) = u u,. Figure I. l(c) shows the free body diagram of the girder, in which f; denotes the inertia force, f, the spring (or the column) force and f, denotes the damping force. The equilibrium equation for the girder is simply


Our structure is linear elastic, having the force-displacement relationship shown in Figure I. l(d). Therefore,f, = ku. The viscous damping force f, is assumed to vary linearly with relative velocity u, f, = cii, Figure l.l(e). The inertia forcef; u,). A super dot ( ' ) denotes the time derivative. Making the is given by tn(u substitutions in Equation 1.1, we get



+ ii,) + cu + ku = 0,

( 1 ..a

Equation 1.3 represents damped vibrations of the structure subjected to the -mu, force. We now use the following basic relationship of structural dynamics; k = t?ro2, and c = 2mol; which with Equation 1.3 becomes

where o is circular frequency of the structure in radians per second and [ is the damping ratio. For free response to be vibratory, ( < 1. For most structures l; is small, say < 0.1, or 10%. We note that the frequency in Hertz (Hz) or in cycles per second (cps)f = 0/2n, and that the period of vibration T = I /f= 2x10, which is in seconds. Equation 1.4 can be solved using standard numerical techniques. As a result we can obtain the time histories of displacement, velocity and acceleration, of the spring and the damping forces. and any other related response time history. See the Appendix. .


Mass M

Lateral Stiffness


(a) One story model


(b) Model subjected to ground motion


(c) Free body diagram

(d) Elastic force-deformation relation

(e) Viscous damping forcevelocity relation

Fig. 1.1 A single-degree-of-freedom model. (Based on Chopra [I].)

1.2 Response spectrum We can solve Equation 1.4 for many single-degree-of-freedom (SDOF) structures having different frequencies, each subjected to the same earthquake ground motion. For each structure we can calculate the absolute maximum value of the response of interest from the corresponding time history. In earthquake response calculations the sign of response is often not considered. For design purposes the maximum positive and negative values are assumed to have equal magnitudes, hence the absolute sign. The curve showing the maximum response versus structural frequency relationship is called the response spectrum.


Time, sec Fig. 1.2 Ground acceleration history of El Centro earthquake (SOOE,1940).

For designing a structure, we are most interested in the maximum spring force

f,,which can be evaluated if the maximum relative displacement u is known. A plot between maximum relative displacement and structural frequency is called the displacement response spectrum. Its ordinates are called spectral displacements, and are denoted by S,(f; (). Depending upon the context, they can also be SD(f ), SD(o),or simply by SD.Let us write denoted by SD(o,0,

Figure 1.2 shows the ground acceleration time history of the El Centro (SOOE, 1940) earthquake. The corresponding displacement response spectrum is shown in Figure 1.3(a). Let us consider the spring force-displacement relationship f, = ku. We have indicated earlier that if the relative displacement u is known, we can find the spring forcef,. Alternatively, if the spring force is known, we can determine the corresponding relative displacement. We can visualize this as a pseudo-static problem shown in Figure 1.4. Now let us think of-&as a pseudo-inertia force, which can be written in terms of the pseudo acceleration a as ma. The relationships, ma = f, = ku, give a = (k/m)u = 0 2 u . The absolute maximum value of a is called spectral acceleration SA.We can easily see

From Equation 1.2 we observe that when cu is small we can write m(ii ii,) -- -ku, or the total acceleration (u u,) 1. (-k/m)u = - 0 2 u . This means,



This pseudo-acceleration response spectrum for the El Centro earthquake is plotted in Figure 1.3(c).

Fig. 1.3 (a) Displacement response spectrum. (b) Velocity response spectra. (c)

Acceleration responsc spectrum for El Centro earthquake (SOOE. 1940);damping ratio, 6 = 0.02.


Fig. 1.4

Pseudo-static problem.

Having defined the response spectra for relative displacement and for pseudo acceleration, we wish to define a response spectrum for velocity. It can be done in more than one way. First, let us define a spectral velocity Svsuch that the kinetic energy associated with it is equal to the maximum strain energy of the spring, . gives (1/2)mS: = ( 1 / 2 ) k ~ iThis

The spectral velocity Sv is really a pseudo velocity because it is not directly related to the actual velocity of the structure. This pseudo-velocity response spectrum for the El Centro earthquake is plotted in Figure 1.3(b). We now have three spectral quantities SD,Sv and SAwhich have units of displacement, velocity and acceleration, respectively. Only spectral displacement SDis directly based on an actual response quantity, the maximum relative displacement. Equations 1.6 and 1.8 give their mutual relationships

Because of this relationship it is possible to read S,, Svand S, from the same logarithmic chart shown in Figure 1.5. This chart is known as the tripartite chart because, for any frequencyJ there are three scales, one each for SD,Svand SA. Now consider the second way of defining a velocity spectrum. We shall denote the new quantity by S;. It is defined as the absolute maximum relative velocity



= max u ( t )



The relative velocity spectrum is shown in Figure 1.3(b) with the dashed lines. The two spectra in the figure are close in the intermediate frequency range; the pseudo velocity spectrum is higher in the high frequency range, and the relative velocity spectrum is higher in the low frequency range. Thus, as a rule, we cannot substitute one spectrum for the other.

For the SDOF structure, the response spectrum quantity of interest is any one of S,, S, or S,. Also, for the classically damped r n u ~ ~ ~ ~ f - f r e e d o m spectra. (MDOF) systems defined in Section 1.4, we need only one o&*drce We shall see in Chapter 5, that we also need S; for nonclassically? systems.

1.3 Characteristics of the earthquake response spectrum Let us observe Figure 1.5 again, which shows the tripartite El 1940) response spectrum, along with the maximum ground displacement, velocity and acceleration values. It is clear that in the low frequency range S, = max I u, I, and in the high frequency range S, = max 1 ii, I . This phenomenon can be easily explained. The low frequency range is characterized by a low value of the spring stiffness k, o = J ( k / m ) . As the spring stiffness becomes smaller and smaller, it progressively ceases to transmit any motion to the mass. In the limit, the total displacement of the mass tends to zero. Relative displacement of the oscillator becomes - u,, or S,, = max I u, I .



Maximum relative displacement can be expressed as: S, = I u I mii,/k,,,I X (dynamic amplification factor). We know that when the oscillator (structural) frequency is sufficiently greater than the dominant frequencies of the input force (mu,), then the dynamic amplication factor = 1.


Therefore, S,


/ mu,/k I ,,, = ( i i , / 0 2 I ,,

or S,

= 02~,, = ( ii,


We can think of the tripartite response spectrum as 'anchored' on the two sides to the maximum ground displacement and acceleration values. In the intermediate frequency range the spectrum has amplified spectral displacement, velocity and acceleration. These observations will be useful in developing design spectra in the next chapter.

1.4 Multi-degree-of-freedom systems Figure 1.6 shows a 3-degree-of-freedom (3-DOF) structure which is a simple example of MDOF systems. The equation of motion for this structure can be derived in a manner similar to that for the SDOF structure we did earlier. For a rigorous derivation the reader is referred to books on structural dynamics [2]. Our example 3-DOF structure has three story masses, nt, , m,,m,, and three story stiffnesses, k,, k,, k,. The three DOF are associated with the lateral (horizontal) displacements of the three masses. The structure deforms under the action of earthquake ground motion, u,. The relative displacement of the structure is given by ur = [u, u2 u,]. The inertia force vector is



0 in2 0


+ ii,

[ii2+iig], in3


(1.1 1)

+ ii,

where U, is a vector of ground acceleration ii,. The vector of spring forces is given by

When damping is absent the equilibrium equation simply becomes F, which can be written as

+ Fs = 0,

In the above equation M is the mass matrix of the structure, K the stiffness matrix, and the vector 1 consists of unit elements. For the 3-DOF structure these matrices are explicitly defined above. For other MDOF structure these matrices

(a) Simple 3-DOF System

(b) Deformed Shape

Fig. 1.6 Example of a multi-degree-of-freedom system.

can be obtained using standard procedures [2]. A more general form of the undamped equation of motion is


The vector U,,defines static structural displacements when the support undergoes a unit displacement in the direction of the earthquake. For the simple structure at hand, it is easy to see that U,becomes 1 as in Equation 1.13. The mode shapes and the frequencies of the structure are obtained by solving the following eigenvalue problem

where w is a natural frequency of the structure. The solution of Equation 1.15 gives N frequencies and the corresponding mode shapes or modal vectors, where N is the number of DOF of the structure. Figure 1.7 shows the mode shapes and frequencies of 3-DOF structure when m ,= m, = m3= m, and k, k, = k, = k. Let us denote the frequency of the ith mode of an N-DOF structure by a, and the modal vector by 9,. The modal vectors have the following orthogonal properties





( P ~ ' M @ ~ = O and @ i r ~ @ j = O f o r i # j . The modal vectors are often 'normalized' such that

in which case, it can be shown that



Mode 1 f = o.0708


Mode 2 Hz


= 0 . 1 9 8 4 HZ

Fig. 1.7 Unnormalized mode shapes and frequencies of a m, = m, m , , k , = k 2 = k,.



Mode 3 0.2a7fi



The response of the structure is represented in terms of a linear superposition of mode shapes

where y, terms are called normal coordinates, and are functions of the time variable t. Substitution of Equation 1.17 in Equation 1.14, premultiplication by i$jT, and the application of the orthogonality conditions from Equation 1.16 gives

in which y, is called the participation factor for mode i, and is given by

Equation 1.18 is similar to Equation 1.4 for the SDOF structure for the undamped case. It is difficult accurately to define the damping matrix for a MDOF structure. Often it is assumed that the damping matrix C has orthogonality properties similar to those of M and K,and that we can define the damping ratio for each mode just as we did for a SDOF structure

Structures that have the idealized damping matrix property given by the above equations are called classically damped. Equation 1.18 is replaced by

In the modal superposition method Equation 1.21 is solved to obtain the time histories of the normal coordinates y,, which with Equation 1.17 give the history of the relative displacement vector U,etc. We shall now use the above concept to apply the response spectrum method to the MDOF structure. The comparison of Equations 1.4 and 1.21 shows, yi(f) = y,u(t), when o = w, and & = 6,. Hence, yimar= y, SD(w,. = y,SD;. Thus, the maximum displacement vector in the ith mode can be written as


Given the displacement vector U,,,,,we can determine the maximum value of any response of interest. Methods of combining maximum response values from various modes, and from three components of earthquake are presented in Chapters 3 and 4, respectively.


I . A.K. Chopra, Dwamics of Struc~ures- A Prirtier, Earthquake Engineering Research Institute, Berkley, California, I98 1 . 2. R.W. Clough and J. Penzien, D7vnarnicsofStructures, McGraw Hill, New York, 1975.

Chapter Z/Design spectrum

2.1 Introduction We have reviewed the basic concepts of dynamic structural analysis in Chapter 1. Only linear elastic behavior has been considered. The purpose was to set the stage for other topics. Indeed, theory and techniques of structural dynamics have reached a stage of advancement such that it is fair to say that any structure that can be mathematically modeled can be analyzed subject to any given transient forcing function, e.g. an earthquake ground motion. The structure may have any given linear or nonlinear constitutive properties. Large displacement effects can also be accounted for. For us the key here is the definition of the earthquake ground motion. If we know the ground motion history, we can analyze the structure and design it. But the earthquake we are talking about has not yet occurred. In many ways the problem of specifying a future earthquake is not very different from specifying any other load for design. The actual live load on a building floor varies a great deal during its lifetime, and it is not uniformly distributed on the floor area. There are at least three idealizations involved in live load specification. First, we idealize the actual floor distribution of furniture, people and other live loads as a uniformly distributed load such that the design quantities of interest-the slab moments-have approximately the same magnitude. Second, we estimate the likely maximum magnitude of this uniformly distributed load during the lifetime of the building. Finally, we design the floor using appropriate load factors, capacity reduction factors or factors of safety. The end product is a slab, which, incidentally, has a relatively definite resistance. The art of specifying the load and the remainder of the design procedure is then a 'recipe', which more than anything else assures a resistance. Therefore, when we are specifying a load, we are really specifying the resistance or the level of resistance in a structure or a structural element. The specification of the earthquake load consists of determining themagnitude and intensity of the design ground shock at a given site, and of somehow converting them into the ground motion parameters. The intensity of the design earthquake is determined from the seismological and geological data concerning earthquakes and their occurrence. It is well to note that available data base is far from adequate and is the major source of uncertainty in earthquake-resistant design. It is then unrealistic in most cases to expect that we can characterize a future ground motion in any detail based on the design intensity and any other limited information available. Approximate procedures have been developed to 11

give the estimate of the peak ground acceleration associated with intensity levels. In some cases the peak ground displacements and velocity can also be approximately estimated. Given the peak ground motion parameters-displacement, velocity and acceleration-techniques have been developed to define smooth spectrum curves, which are called design spectra. The main difference between the response spectrum presented in the previous chapter and the design spectrum we are discussing now is that the former represents the response to an actual earthquake and the latter defines the level of seismic resistance we are to design for. Just as in the example of the live load, the design spectrum idealizes the real phenomenon to fit it into a design recipe.

2.2 'Average1 elastic spectra Biot [I .2] and Housner[3] were among the first researchers who recognized the potential of the response spectrum as an earthquake-resistant design tool. Biot[l.2] developed a mechanical analyzer to evaluate experimentally the response of a single-degree-of-freedom system subjected to recorded earthquake acceleration time histories. For design purposes he suggested smoothing the response spectra obtained from actual records. The mechanical analyzer Biot used was practically undamped, although he recognized that the damplng will lower the peaks of the response spectra. Later, Housner[4] obtained design spectra by averaging and smoothing the response spectra from eight ground motion records, two each from four earthquakes, viz., El Centro (1934), El Centro (1 WO), Olympia (1949), Tehachapi (1952). Housner's spectra for several damping values are shown in Figure 2.1. These were the first spectra used for the seismic design of structures. The spectra shown in Figure 2.1 have been scaled for 0.125 g peak ground acceleration. One could scale them to any other peak ground acceleration consistent with the design intensity of earthquakes at the site. Newmark and coworkers[5,6] studied the response spectra of a wide variety of ground motions, ranging from simple pulses of displacement, velocity or acceleration of ground through more complex motions such as those arising from nuclear blast detonations and for a variety of earthquakes as taken from available strong motion records. They observed that the general shape of a smoothed response spectrum looks like that shown in Figure 2.2 on a logarithmic tripartite graph. As we observed in Chapter 1, in the low frequency range, the special displacement SD= maximum ground displacement d; and in the high frequency range, the spectral acceleration S, = maximum ground acceleration a. The intermediate frequency range can be divided into five regions: ( I ) an amplified velocity region in the middle which is flanked by (2) amplified displacement and (3) acceleration regions, and two regions of transition, (4) from the maximum ground displacement to the amplified spectral displacement and (5) from the amplified spectral acceleration to the maximum ground acceleration. In an earlier publication, Blume, Newmark and Corning[7] had suggested the following

DESIGN S P E C T R U M / 1 3








Period (sec)

Period (set) Fig. 2.1 (a) Housner velocity design spectra. (b) Housner acceleration design spectra, peak ground acceleration = 0.125 g [4l.


Frequency, f (Log scale) Fig. 2.2 General shape of a smoothed response spectrum.

factors for amplified spectral displacement, velocity and acceleration: respectively 1.0, 1.5, 2.0 for 5- 10% damping ratio; and 2, 3 , 4 for < 2% damping ratio. The elastic spectra of the type we are discussing here are primarily used for critical facilities like nuclear power plants. The United States Atomic Energy Commission sponsored two comprehensive studies in the early seventies, one conducted by Mohraz, Hall and Newmarkt81 for Newmark Consulting Engineering Services and the other by Blume, Sharpe and Dalal[9]. In the Blume study [9] two components of horizontal motion for sixteen earthquakes, and one component for an additional earthquake, a total of thirty-three different records were considered. In the Newmark study[8] fourteen earthquakes were considered with two components of the horizontal motion and one component of the vertical motion for each earthquake. The two studies were conducted independently and there are differences in their details. The conclusions from the two studies were quite similar, however. Therefore, we will present the results of the Newmark study[8] only. In order to 'average' the response spectra from various ground motion records they need to be normalized using a ground motion parameter, maximum ground displacement, velocity or acceleration. It was found[8] that the normalization to the maximum ground acceleration gave a standard deviation that increased quite uniformly from high frequencies to low frequencies. Normalization to maximum ground displacement showed a standard deviation which increased from low to high frequencies. Further, the normalization to maximum ground velocity showed a nearly constant standard deviation over the whole range of frequencies. The smallest standard deviation was obtained in each region when the normaliza-


tion was made to the particular ground motion parameter for which the response spectrum curve was most nearly parallel to the coordinate. In the Newmark study[8], the spectral ordinates were assumed to have a normal distribution; the data would also fit the log-normal distribution[9]. The design spectrum can be obtained from the maximum ground displacement, velocity and acceleration values, if the respective amplification factors are known. These amplification factors were obtained for the mean spectrum (50% probability level) and for the mean plus one standard deviation spectrum (84.1% probability level). Table 2.1 summarizes these amplification factors for the horizontal components of earthquake for four damping values. Values simiIar to those in Table 2.1 for 84.1% probability level have been adopted by a recent ASCE standard[lO]. In the Newmark study[8], it is recommended that the transition from amplified acceleration to ground acceleration begin at 6 Hz for all damping values and end at 40, 30, 20, 20 Hz, respectively for damping ratios of 0.5%, 2.0%, 5.0% and 10.0%. In the ASCE Standard[lO], this transition occurs between 9 and 33 Hz for all damping values. Corresponding to a Ig maximum ground acceleration, it was found[8] that the maximum ground displacement and velocity were approximately 36 in and 48 in sec-I, respectively, for alluvium soil, and 12 in and 28 in sec-' for rock. For both types of soiis, we have ad/u2 2: 6, the value recommended by the ASCE Standard[lO]. Figure 2.3 shows the 'Newmark spectrum' for Ig maximum ground acceleration on an alluvium soil. The joint Newmark-Blume recornmendations[l l ] were later adopted in a modified form by the United States Atomic Energy Commission (now the US Nuclear Regulatory Commission-USNRC)[12]. Figure 2.4 shows the USNRC spectra for Ig ground acceleration. As can be seen, there are no major differences between Figures 2.3 and 2.4. The recommendations for the vertical component of

Table 2.1 Amplification factors for Newmark spectrum[8]

Damping ratio (%) Spectral quantity

Probability level (%)




SD = fact0r.d. Sv = factor.^, SA = fact0r.a. d. v, a = maximum ground displacement, velocity, acceleration, respectively. adlv2 1. 6.


Frequency, Hz

Fig. 2.3 Newmark design spectra for alluvium soil; maximum ground acceleration = I g.

earthquake vary much more, and a complete discussion is beyond the scope of this book. The ASCE Standard[lOJ recommends that the design spectra for the vertical component be obtained by multiplying the corresponding spectra for the horizontal component by a factor of 2/3.

2.3 Site-dependent spectra The spectral amplification factors presented in the previous section are based on the analysis of several earthquake ground motions without particular consideration of the site conditions. The only consideration of site has been that for maximum ground motion parameters suggested by Mohraz, Hall and Newmark[8], d = 36 in and u = 48 in sec-' for alluvium site; and d = 12 in, 28 in sec-' for rock site; for I g maximum ground acceleration. One would expect that the site conditions influence the frequency content in the ground motion, and therefore, the spectral amplification factors would depend upon them too. Mohraz. Hall and Newmark indicated that the spectrum for an alluvium site is considerably different from that for a competent rock site. Since only a few accelerograms from stations on rock site were considered in their study[8], no conclusive recommendations were made.

DESIGN S P E C T R U M 1 1 7











Frequency. Hz

Fig. 2.4 United States Atomic Energy Commission design spectra [12].

A major problem associated with evaluating these site-dependent spectra is in

the description of the site itself. One possible method is to classify the recording stations according to their shear wave velocity. For most stations the estimates of shear wave velocity are not available, nor available are details of any other soil properties at the stations. Researchers have, therefore, used the limited site information and their experience, and have subjectively classified the recording stations. As we would expect, different researchers have different classifications. Studies that have been performed do show definite trends, and in that sense they are very valuable. Hayashi, Tsuchida and Kurata[l3] performed a study in which they averaged the normalized response spectra for sixty-one accelerograms obtained from thirty-eight earthquakes in Japan including many with peak accelerations in the relatively low range of 0.02-0.05 g. The spectra were averaged in three groups. Group A was considered to be associated with very dense sands and gravels, Group B with soils of intermediate characteristics and Group C with extremely

loose soils. As expected, they found that the soil conditions affect spectra substantially. Although the authors suggested that the spectral shapes be considered preliminary in view of the relatively low maximum accelerations associated with the earthquake records and the limited data on some of the subsoil conditions at the recording stations, their results were later found to be in substantial agreement with those obtained by Seed, Ugas and Lysmer[l4]. A limited study of the influence of local site conditions on spectral shapes for Japanese earthquakes was presented by Kuribayashi el al. [ I 51. All the studies on the site dependence of the spectra cited above were performed for the horizontal components of the ground motions. A recent comprehensive study by Seed, Ugas and Lysmer[l4] was also performed on the horizontal components only. They considered four site conditions: rock, stiff soils less than about 150 ft deep, deep cohesionless soils with depths greater than about 250 ft, soil deposits consisting of soft to medium stiff clays with associated strata of sands or gravels. Their study was based on one hundred and four records each with maximum acceleration >0.05 g. The mean plus one standard deviation-84.1 percentile spectra obtained by Seed, Ugas and Lysmer[l4] for the four site conditions, 5% damping ratio are shown in Figure 2.5 along with the corresponding AEC spectrum[l2]. All the spectra are normalized for 1 g maximum ground acceleration. There are wide differences in the spectral shapes for the four sites for periods greater than roughly 0.4 sec. In these period ranges, the sites on softer soils (deep cohesionless soil and soft to medium clays and sands) have much higher spectral values than those on stiffer soils (rock and stiff soil). The AEC spectrum has the best agreement with Seed, Ugas and Lysmer's spectrum on the stiff soil in the 0.5-3 sec period (0.3-2 Hz frequency) range. The agreement between the AEC spectrum and Seed, Ugas and Lysmer's spectra for all the four site conditions is generally good in 0.1-0.5 sec (2- 10 Hz frequency) range. For periods < 0.1 sec, or frequencies > 10 Hz, the AEC spectrum significantly overestimates the spectral values for the soft soils. Another recent significant study is due to Mohraz[16]. His study included the vertical components of earthquake-in addition to the usual horizontal components. Further, he calculated and recommended spectral values for several damping ratios. He studied the effects of geological conditions on the spectra, and also on the ground motion parameters, such as peak ground acceleration, velocity and displacement. He considered the two common site conditions, alluvium and rock, and two intermediate site conditions-deposits of < 30 ft of alluvium, and deposits of approximately 30-200 ft of alluvium, both underlain by rock deposits. One reason the latter two categories were selected was that substantial earthquake records from stations located on such deposits were available. The sites labeled alluvium, were those which did not have a specified depth, and may have a depth less than or greater than 200 ft. Mohraz analyzed one hundred and six records from forty-six stations in sixteen seismic events.


Soft to Medium Clay and Sand eep Cohesionless Soik (>250W Stiff Soil Conditions k 1 5 0 f t )

Regulatory Guide








Period, Seconds Fig. 2.5 Seed, Ugas and Lysmer site-dependent spectra and Atomic Energy Commission spectrum; mean plus one standard deviation (84.1 percentile), peak ground acceleration = 1 g, damping ratio = 0.05 [14].

Mohraz had three components for each earthquake. The peak accelerations for the three components are ordered from the largest to the smallest as follows: the larger horizontal a,,, the smaller horizontal a,, and the vertical a,. The mean value of the ratios aSH/aLH and a,/aLH for all the site conditions are 0.83 and 0.48, respectively. The corresponding 84.1 percentile values are 0.98 and 0.65. That means that at the 84.1 percentile level the maximum ground acceleration for the smaller horizontal component is almost equal to the maximum horizontal acceleration for the larger horizontal acceleration, and that the maximum vertical acceleration is almost 213 of the latter. This is consistent with the assumption commonly made for the design of critical facilities[lO]. Mohraz performed a detailed statistical study on the ratios of the ground motion parameters, u/a and ad/v2. When the ratios are known, we can calculate the values of u and d; the value of a is commonly given based on seismological considerations. Table 2.2 lists the mean values of the ratios v/a and ad/v2, along with the corresponding values of u and d for various site conditions for a 1 g maximum ground acceleration. For illustration here, the larger horizontal

Table 2.2 Ground motion parameters[ 161 Larger horizontal component For a- l g

Mean ratios vl0

Site condition

Vertical component

adlv2 v

(in sec-'g-')

For a- l g

Mean ratios



(in sec-') (in)

(in scc-'g-')

ad/u2 v


Based on Mohraz[l6] Rock









Less than 30 ft alluvium underlain by rocks









30-200 ft Alluvium underlain by rocks


















6 6

28 48

12 36





Based on Mohraz, Hall and Newmark(81 Rock Alluvium

28 48



component and the vertical component are included in Table 2.2. The table indicates that the v / a ratios for rock are substantially lower than those for alluvium. These ratios for the two alluvium layers underlain by rock are between those for rock and alluvium. The v / a ratios for the larger horizontal components are the same as that for the vertical component, except in the case of rock, when they are close. The ad/v2ratios given in Table 2.2 indicate that the ratios for alluvium are smaller than those for rock and alluvium layers underlain by rocks. Since ad/u2 is a measure of the width of the spectra, the ratios indicate that the average spectrum for a rock site is flatter than for alluvium site or for a site with alluvium layers underlain by rock. The values of vla and adlv2 ratios, and of v and suggested earlier by Mohraz, Hall and Newmark(81 for the horizontal component only are also given in Table 2.2 for comparison. There are minor differences between the old and the new v/a values. A relatively greater change occurs in the value of ad/v2for the alluvium site (6-4.3) and in the corresponding value of d (36-28.9 in). We note that Mohraz[I 61 also gives the median and 84.1 percentile values, which d o show dispersion in these values. For most practical applications we believe that using the mean values of the ratios should be adequate. Mohraz(l61 showed that his average spectrum for the rock site, 5% damping ratio compared well with the corresponding spectrum given by Seed, Ugas and

DESIGN S P E C T R U M / 2 1

4 -




Lesa than 30 ft. Alluvium on Rock

-. ,.

30-200 ft. Alluvium on Rock

, , Rock


0 0








Period, Seconds Fig. 2.6 Mohraz average site dependent spectra; peak ground acceleration ration = 0.02 [16].



I g, damping

Lysmer[l4]. Mohraz's average spectra normalized to 1 g for the four sites are shown in Figure 2.6 for 2% damping ratio. We observe from Figure 2.6 that the acceleration amplification for the alluvium deposits extends over a larger frequency region than the amplification for other site categories. Further, the maximum spectral accelerations for the two sites with alluvium layers underlain by rock is greater than the maximum spectral acceleration for either rock site or the alluvium site. For periods ~ 0 . 2sec (frequencies > 5 Hz), the spectral ordinates for alluvium sites are less than those for the other three site types studied. In the periods >0.5 sec (frequencies < 2 Hz), the spectral ordinates for the alluvium site are greater than those for the other three. The alluvium spectral values are approximately 2.5-3 times the rock spectral values in the 1.5-3 sec period range (frequencies, 0.3-0.7 Hz). Mohraz [I 61 has presented comprehensive statistics of displacement, velocity and acceleration amplification factors for all the four site types, damping ratios 0-20%, and the three components of earthquake. Amplification factors are given for the mean (50% probability level) and the mean plus one standard deviation (84.1% probability level) spectral values. Amplification factors for the larger horizontal component for the alluvium site only are reproduced in Table 2.3, for three damping ratios (2%, 5% and 10%) and for 50% and 84.1% probability levels.

Table 2.3 Amplification factors for larger horizontal component for alluvium site

suggested by Mohraz(l61 Damping ratio (%) Spectral quantity

Probability level (%)




= factor-v, S , = fact0r.a. maximum ground displacement, velocity, acceleration, respectively; Table 2.2.

S,,= fact0r.d. S,, d, v , a


For intermediate damping ratios, Mohraz[l6] suggests double logarithmic interpolation. Mohraz has given amplification factors for the other three site conditions also. For brevity, we are not reproducing that information here. Instead, in Table 2.4 are given the site design spectra coefficients, also reproduced from Mohraz[l6], which can be used to obtain the spectral bounds for the other three site conditions from those for the alluvium site. Mohraz recommends the same coefficients for the two alluvium sites underlain by rock deposits because the coefficients for these two categories do not vary significantly from each other. Since the number of available records for these two site types is not as large as that for either the rock or the alluvium deposits, the recommended coefficients are on the conservative side. \

Table 2.4 Site design spectra coefficients[l6]

Coefficients Site category








Less than 30 ft alluvium underlain by rock



I .20

30-200 ft Alluvium underlain by rock




Design spectrum value at the site = site coefficient X design spectrum value at the alluvium site.

DESIGN S P E C T R U M / 2 3

Frequency, Hz Fig. 2.7 Mohraz and Newmark site dependent spectra; peak ground acceleration

= 1


damping ratio = 0.02.

Given the maximum ground acceleration, there is sufficient information in Tables 2.2,2.3 and 2.4 to obtain 50%or 84.1% probability level design spectra for the larger horizontal component. Although the Mohraz study shows that the other horizontal component has slightly less maximum ground acceleration than does the larger horizontal component, it is a common practice to assume that the design spectra for the two horizontal components are the same. Mohraz[l6] does give much detail about the vertical component of the earthquake. However, at the end he derives the vertical design spectral ordinates as 2/3 of the ordinates of the horizontal spectrum, which is consistent with the present practice. The Mohraz horizontal spectra for 2% damping ratio, 84.1 percentile level are shown in Figure 2.7 for the three site conditions; (based on Table 2.4, the two alluvium sites underlain by rock are combined into one site type). Also shown in Figure 2.7 are Newmark's spectra[8] for alluvium and rock sites.

2.4 Design spectrum for inelastic systems The basis of applying the response spectrum method to multi-degree-of-freedom systems is the modal superposition method; see Chapter 1. Strictly speaking,

therefore, the method cannot be applied to inelastic multi-degree-of-freedom systems because the superposition is no longer valid. No such difficulty, however, exists when a single degree-of-freedom is under consideration. In that case the response spectrum simply represents the maximum value of the relative displacement-or of any other quantity of interest. The maximum value can be evaluated whether the system is linear or nonlinear. In this section we shall present the inelastic spectrum for the single-degree-offreedom systems. When the major response of a structure, such as a tall building, comes from the fundamental mode, then we can consider the structure to be a pseudo single-degree-of-freedom system and make use of the inelastic spectrum for evaluating the required resistance of the structural members. The inelastic spectrum is sometimes also used for calculating response in higher modes. The accuracy of such an approach is questionable. It can be justified because the contribution of higher modes is relatively small, and the error in the evaluation of higher mode response would not introduce significant error in the overall response of the structure. The question of inelastic response of multi-degree-offreedom systems is a complex one, and it continues to be a topic of active research. A full discussion on the topic is beyond the scope of the present work. We do note that it is an important topic-the majority of buildings and many other structures are designed based on the assumption of significant inelastic response in case of a severe earthquake. The simplest inelastic material is elastic-perfectly plastic with equal yield values in tension and compression. For a single-degree-of-freedom system, the ductility factor or ductility (p) is defined as the ratio of the maximum . linear elastic analysis deformation urnto the yield deformation u,, p = u r n / u yIn we assume that maximum deformation remains below u , . The member is designed such that the analytically calculated u is less than or equal to u , ; or that the analytically calculated member force or stress o is less than or equal to the yield force or stress o, corresponding to u,. If the system is capable of safely undergoing inelastic deformation, it is economical to design it such that the maximum allowable deformation is achieved under the given earthquake. For a given material and structural system, the permissible ductility factor p can be judged to be known. The objective of the calculation is to evaluate u , such that u, is achieved under the given earthquake. In Equation 1.1 ,.A fD + -f, = 0,1; is still m(u u,) and fD remains cu. Due to inelasticity now, S, = ku when ) u I G u , , and& = + o, when ) u 12 u,. After one or more plastic excursions these conditions are appropriately modified. The solution of the nonlinear equation is straightforward, although more involved than the solution of a linear equation; see Appendix. The response spectrum consists of the response from many single-degrees-of-freedom of systems with varying frequencies; the damping is kept constant for each response spectrum curve. Now we have another variable, the ductility ratio p. For the ductility ratio



DESIGN S P E C T R U M / 2 5

of unity, urn= u,, and we have the elastic response spectrum curve. The inelasticity is introduced when p > 1. Again, for each response spectrum curve a constant value of p is assumed. For a single-degree-of-freedom system of given frequency and damping, the solution process consists of assuming a value of u,, and integrating the nonlinear equation of motion numerically for the earthquake ground motion history which gives u,. For the assumed value of u,, the calculated urnis not likely to give the ductility factor urn/uyequal to the desired p. For each point, therefore, several solutions have to be performed, each with a different value of u,. When the calculated urn/uyis sufficiently close to the desired p, the iteration stops. The elastic response spectrum is based on maximum relative displacement, which is also a measure of the maximum spring force. We recall, the relationship S, = oZSD is obtained on the basis that the pseudo-inertia force given by S, is equal to the spring force given by SD.For an inelastic single-degree-of-freedom system the measure of spring force is u,. Since u, and urnare so conveniently related, urn= pu,, we can use either of the two displacements for drawing the response spectrum, as long as we know which one it is in a given case. The spectrum based on urncan be called the maximum displacement spectrum, and that based on u , , the yield spectrum. Early studies on the inelastic response spectrum were made by Newmark and coworkers [5,6,17]. They reached the following conclusions: 1. For low frequency systems, the maximum displacement for the inelastic system (urn)is the same as for an elastic system having the same frequency. 2. For intermediate frequency systems, the total energy absorbed by the spring is the same for the inelastic system as for an elastic system having the same frequency. 3. For high frequency systems, the force in the spring is the same for the inelastic system as for an elastic system having the same frequency. Let us denote the elastic spectral values by s E , and the corresponding maximum displacement and yield spectra values by SMand SY,respectively. The above conclusions give the following relationships: Low frequency range,

SM= s E ,sY = SE/p;

Intermediate frequency range, SM= High frequency range,

P p , SY J ( ~ P- 1)

SM= p s E ,sY = sE.




- I) (2.1)

Note, in all the frequency ranges, sM = psY. These recommendations are applied to the Newmark type elastic design spectrum in Figure 2.8. The symbols D, V, A, A, refer to the bounds of the elastic

Frequency (Log Scale)

Fig. 2.8 Newmark inelastic response spectra.

spectrum. A, represents the maximum ground acceleration. Superscripts M and Y are used to denote the corresponding maximum displacement and yield spectral values. The elastic spectrum bounds D and Vare covered by the small frequency range. The corresponding D Yand V' values are obtained by dividing D and Vby p. The value ofA is obtained by dividing A by J(2p - 1). A: remains the same as A,. The D and Vspectral lines, and the DY,A' spectral lines intersect at the same key frequency; the key frequency at the intersection of VYand A 'is in general different from that at the intersection of V and A. Usually, we begin the transition from A to A, at the same frequency at which the transition from A to A, begins. Having obtained the yield spectrum, the evaluation of the maximum displacement spectrum is straight forward. We simply multiply the yield spectrum ordinates by the ductility factor to do so. In the resulting maximum displacement spectrum, we note D~ = D, vM= V. Riddell and Newmark11 81 performed a relatively detailed study on the topic of the inelastic spectrum. They found that the factors used to modify an elastic spectrum into an inelastic spectrum are functions of the damping ratio and of the type of the material force-deformation relationship. Overall, however, they confirmed the conclusions of the earlier studies[5,6,17]. They evaluated inelastic spectra for elastoplastic, bilinear and stiffness degrading systems. They concluded that the ordinates of the average inelastic spectra for the three material models did not differ significantly. They also found that the spectrum for the elastoplastic material was always slightly conservative as compared with those for the other two materials. That leaves the effect of damping values. One may use the more refined modification factors given by Riddell and Newmark[lS].








DESIGN S P E C T R U M 1 2 7

However, the damping independent modification factors reported here based on the earlier work[5,6,17], appear to be adequate for most practical applications. The dependence of the inelastic spectra on the sites was studied by Elghadamsi and Mohraz[l9]. As explained above, the inelastic spectrum for a given ductility level has to be obtained iteratively. This procedure is computationally inefficient. Elghadamsi and Mohraz computed the maximum displacement spectra for a given yield displacement, u,. This eliminated the iterations. The authors found that their procedure required approximately 5-10% of the computational time needed for the ductility based inelastic response spectrum calculations. This permitted them to consider a relatively large ensemble of earthquake ground motions, fifty records on alluvium sites and twenty-six on rock sites. The maximum displacement spectra for fixed yield displacement values, in effect, represent the ductility demand curves. These curves can be easily converted into the constant ductility inelastic spectra. Elghadamsi and Mohraz[l9] used four material models: elastoplastic, bilinear, four-parameter Nadai and a new stiffness degrading model. They found that the inelastic design spectrum computed from the elastoplastic model can be used conservatively in most cases to estimate the design spectrum from the other three models.

2.5 Comments An effort has been made in this Chapter to summarize some of the available information on design spectra in a simplified form. For many readers the information presented here may be adequate. For others, especially, those involved in the design of major facilities and those interested in pursuing the topic for research, there is a vast growing body of literature available. A philosophical and historical perspective is given by Newmark and Ha11[20] and Housner and Jennings[2 11 in two EERI monographs. Broadly classified spectral shapes presented in this Chapter should serve a useful purpose when more detailed and precise data f0r.a site can not be obtained. It is not uncommon in many cases that the only motion parameter that is specified is the peak ground acceleration. The use of peak ground acceleration in conjunction with a standard spectral shape for a site in the vicinity of a fault can grossly overstate the response values in the frequency range of interest. It is much more desirable somehow to evaluate and use the three major motion parameters: the peak acceleration, velocity and displacement. For sites like those in Mexico City, specific knowledge about the local conditions is very important. Housner and Jennings recommend[2 I]: 'A much better method of describing the ground motion simply would be to compare it to a known accelerogram, such as recorded in Taft, California in 1952, or to a synthesized accelerogram. The decription could thus be phrased as: 1.5 times as intense as Taft 1952, with duration of strong shaking 1.2 times as long and with similar frequencies of motion.'

References 1. M.A. Biot, A Mechanical Analyzer for the Prediction of Earthquake Stresses, Bulletin of the Seismological Societv of America, Vol. 31, 194 1, pp. 15 1 17 1. 2. M.A. Biot, Analytical and Experimental Methods in Engineering Seismology, Proceedings, ASCE, Vol. 68, 1942, pp. 49-69. 3. G.W. Housner, An Investigation of the Effects of Earthquakes on Buildings, Ph.D. Thesis, California Institute of Technology, Pasadena, California, 1941. 4. G.W. Housner, Behavior of Structures During Earthquakes, Journal of Engineering Mechanics Division, ASCE, Vol. 85, No. EM4, 1959, pp. 109- 129. 5. N.M. Newmark and AS. Veletsos, Design Procedures for Shock Isolation Systems of Underground Protective Structures, Vol. 111, Response Spectra of Single-Degree-ofFreedom Elastic and Inelastic Systems, Report for Air Force Weapons Laboratory, by Newmark, Hansen and Associates, R T D TDR 63-3096, June 1964. 6. A.S. Veletsos, N.M. Newmark and C.V. Chelapati, Deformation Spectra for Elastic and Elasto-Plastic Systems Subjected to Ground Shock and Earthquake Motions,. Proceedings. Third U,,'orldConference on Earthquake Engineering, New Zealand, 1965. 7. J.A. Blume, N.M. Newmark and L.H. Corning, Design ofMuhi-Story ReinJorced Concrete Buildingsfor Earthquake Motions, Portland Cement Association, Chicago. Illinois, 1961. 8. B. Mohraz, W.J. Hall and N.M. Newmark, A Study of Vertical and Horizontal Earthquake Spectra, Nathan M. Newmark Consulting Engineering Services, Urbana, Illinois; AEC Report No. WASH-1255, 1972. 9. J.A. Blume, R.L. Sharpe and J.S. Dalal, Recommendations for Shape of Earthquake Response Spectra, John A. Blume and Associates, San Francisco, California; AEC Report, No. 1254, 1972. 10. American Society of Civil Engineering, Standard for the Seismic Analysis ofSafey Related Nuclear Structures, September 1986. I I . N.M. Newmark, J.A. Blume and K.K. Kapur, Seismic Design Criteria for Nuclear Power Plants. Journal ofthe Power Division, ASCE, Vol. 99, 1973, pp. 287-303. 12. United States Atomic Energy Commission. Design Response Spectra for Seismic Design of Nuclear Power Plants, Regulatory Guide, No. 1.60, 1973. 13. S.H. Hayashi. H. Tsuchida and E. Kurata, Average Response Spectra for Various Subsoil Conditions. Third Joint Meeting. L!S. Japan Panel on Wind and Seismic Efects, UJNR, Tokyo, May 197 1. 14. H.B. Seed, C. Ugas and J. Lysmer, Site-Dependent Spectra for Earthquake-Resistant Design, Bulletin ofthe Seismological Society of America, Vol. 66, No. 1. February 1976, pp. 22 1-243. 15. E. Kuribayashi, T. Iwasaki, Y. lida and K. Tuji, Effects of Seismic and Subsoil Conditions on Earthquake Response Spectra, Proceedings, International Conferenceon Microzonation, Seattle, Washington, 1972, pp. 499-5 12. 16. B. Mohraz, A Study of Earthquake Response Spectra for Different Geological Conditions, Bulletin ofthc Seismological Socie?rl ofAmerica, Vol. 66, No. 3, June 1976, pp. 9 15-935. 17. A.S. Veletsos and N.M. Newmark, Effect of Inelastic Behavior on the Response of Simple Systems to Earthquake Motions, Proceedings, Second World Conference on Earthquake Engineering, Vol. 11, 1960. 18. R. Riddell and N.M. Newmark, Statistical Analysis of the Response of Nonlinear Systems Subjected to Earthquakes, Strrrcrural Rcsearch Series, No. 468, Department of Civil Engineering. University of Illinois at Urbana-Champaign, Urbana, Illinois. August 1974. 19. F.E. Elghadamsi and B. Mohraz, Site-Dependent Inelastic Earthquake Spectra, Technical Report, Civil and Mechanical Engineering Department, Southern Methodist University, Dallas, Texas. June 1983.



20. N.M. Newrnark and W.J. Hall, Earthquake Spectra and Design (Engineering Monograph on Earthquake Criteria, Structural Design, and Strong Motion Records, M.S. Agbabian, Coordinating Editor), Earthquake Engineering Research Institute, Berkeley, California, 1982. 2 1. G.W. Housner and P.C. Jennings, Earthquake Design Criteria (Engineering Monograph on Earthquake Criteria, Structural Design, and Strong Motion Records, MS. Agbabian. Coordinating Editor,), Earthquake Engineering Research Institute, Berkeley. California, 1982.

Chapter 3/Combination of modal responses

3.1 Introduction The equation of motion for an N-degree of freedom system was presented in Chapter I, and is rewritten below

where M , Cand K are mass, damping and stiffness matrices, respectively; U is the relative displacement vector; Ubis the static displacement vector when the base of the structure displaces by unity in the direction of the earthquake; u, is the earthquake ground acceleration; and a super dot (.) represents the time derivative. The structure has N-orthogonal modal vectors $,, i = 1 - N. For the present treatment, we assume that the modal vectors have been scaled such that $, M $, = 1. Also, we recall from Chapter 1, $, K $, = of and $i C q i = 20,C,,, where o,is the circular frequency in radians per second, and is'the damping ratio, both for mode i. In the modal superposition method we use the following transformation (or modal superposition equation)


in which y, is called the normal coordinate. Substitution of Equation 3.2 in Equation 3.1, pre-multiplication by +,*, and use of appropriate orthogonality conditions gives

The term yi is called the model participation factor. If SD(w)represents the displacement response spectrum, and we denote the spectral displacement by S,,,then by definition (Chapter l),

Also, we can write

Equation 3.5 gives the maximum relative displacement for each mode. The superposition equation, Equation 3.2, applies only when we know the time histories of all the modal displacement vectors in all the modes. Equation 3.5, however, does not provide that information. In general, it is unlikely that the


maximum values of U;in different modes would occur at the same time. How should we then combine these maximum modal vaiues? Given the modal displacement vector U;, we can evaluate any other response of interest in the same mode, R,. The vector Ui,,, gives Rim,,. The problem with combining various modal U;,,,stated above also applies to the response R,,,,. For brevity, we shall drop the subscript max. From now on, the term R, would denote the maximum value of the response in mode i. It is obvious that the upper bound of the combined response is given by the absolute sum of the modal values

Goodman, Rosenblueth and Newmark[l] showed that the probable maximum value of response is approximately equal to the square root of the sum of the squares'(SRSS) of modal values

Published in 1953, the Goodman-Rosenblueth-Newmark rule, known as the SRSS rule is still used quite widely. There are circumstances, presented in subsequent sections, in which this rule must be modified. For more early research on this topic, also see Jennings and Newmark (1960)[2], Merchant and Hudson (1 962)[3], Clough (1962)[4], and Newmark et al. (1965)[5].

3.2 Modes with closely spaced frequencies One of the exceptions for the SRSS combination rule (Equation 3.7) arises when the responses to be combined are from modes with closely spaced frequencies. An obvious situation is when frequencies and dampings of two modes are identical. In this case the response histories of the two modes are in-phase. The maximum values in the two modes do occur simultaneously, and they should be combined algebraically. We are already denoting the maximum value of the response by R. Let us denote the response history by R(t). The combination equation in the time domain is

Let us define the standard deviation of the response as follows:

in which td is the duration of the motion. If we assume the earthquake to be a stationary ergodic process, we can write the maximum responses [6] as R


qo, Ri = qioi,


where q and q, are the peak factors. These peak factors are a function of the frequency, and would normally vary from mode to mode and for the combined response. However, since we are primarily interested in modal responses with close frequencies, we can make a simplifying assumption, q = q , for all values of i. Equations 3.8 and 3.9 give

in which q, is called the modal correlation coefficient, and is defined by

Now, Equations 3.10 and 3.12 give

The alternative forms of Equation 3.13 are

In the first of Equation 3.14 we have E, = 1 for i = j; and in the second one we take into account the symmetry property, ciJ = E,~. Equation 3.1 3 or 3.14 is popularly known as the double sum equation [7]. In the correct form of the equation all the sums are algebraic. The US Nuclear Regulatory Commission incorrectly uses an absolute sign in .front of the double summation 181. Consider two modes with equal frequencies and damping values. ~ I, and Equation 3.13 gives Then, E , =


As indicated earlier, the double sum rule correctly gives the combined response as the algebraic sum of the two modal response values. On the other hand, if the two modes had sufficiently separated frequencies, E , , 2: 0,and we would get

which is the SRSS rule. In general, the value of E,, varies between zero and unity. For a given earthquake ground motion the value of E, can be evaluated numerically, in accordance with Equation 3.12. We should make two comments here. First, the double sum rule, or any combination rule for the response spectrum method, is an approximate rule. Even when the value of E,, is evaluated 'exactly' from Equation 3.12, the combined responses would be approximate. When the calculations are performed on several earthquake ground motions, the combination rule would, on the average, give a reasonably accurate estimate of the combined response. The second comment is about the response spectrum method, and is related to the first comment. If the objective were to obtain the response values for a known ground motion record, the appropriate technique would be one of the time-history integration methods, e.g. the modal superposition method. The most appropriate application of the response spectrum method is to design problems for which the future earthquake is not known. For this purpose, we not only need 'average' spectral shapes which are presented in Chapter 2, we also need representative values of the modal correlation coefficient E,,, which on average weuld give sufficiently accurate combined response values. Rosenblueth and Elorduy [9] assumed the earthquake ground motion to be a finite segment of white noise, and assumed the response to be damped periodicof the form e-C"' sinoDt. Based on their work, the correlation coefficient can be written as

in which o, and o, are the circular frequencies of the two modes in radians per second;.. o,,_.grid oD, are the corresponding damped frequencies, "\ r'i;,, = J(l - Sf) o,/, wDJ= J(l - (j) o,; and (:and 6; are the equivalent damping r z t o s wTkh account for the reduction in the response due to the finite nature of the white noise segment.

where s is the effective duration of the white noise segment. The duration s is not the total duration of the ground motion. It is not clear how exactly to evaluate it. Villaverde [ lo] obtained numerically values of s for several ground motions by exploiting its relationship with the expected value of pseudo velocities at different damping values. However, he did not suggest any method of evaluating s for a given response spectrum. We note from Equation 3.16 that the value of effective duration has a significant contribution in the lower frequency range only. In the higher For the lower frequency range, the effective duration of frequency range (;fi the earthquake may be represented by the duration of the strong motion. A measure of the effective duration of the strong motion can be obtained from the Husid plot[I 11, which is the graphical representation of the following function



iii(f) dl

H(t) =

[iii(t) dl


in which t, is the final value oft. By definition, 0 S H ( t ) I:1. The Husid plot for El Centro (SOOE, 1940) ground motion is shown in Figure 3.1. The function H(t) builds slowly initially because of the weak motion at the early phase of ground shaking. In the intermediate duration, the H ( t ) builds rapidly. In the final phase, very little seismic input is developed. It is clear from Figure 3.1 that the intermediate portion of the Husid plot comprises the significant strong motion contribution. For definitiveness, but arbitrarily, the first 5% and the last 5% are


1, = 1.67















Time. sec


Fig. 3.1 Husid plot for El Centro earthquake (SOOE, 1940). (Based on Nau and Hall[l I]).


deleted from the plot. The remaining 90% is defined as the significant strong motion portion as shown in Figure 3.1. This duration for El Centro record is 24.5 sec. In using Equations 3.15,3.16 in conjunction with a design spectrum, the value of the duration s should be specified. By substituting the expressions for and in Equation 3.15 from Equation 3.16, we obtain

r; r'

To avoid the estimation of the effective duration s, Gupta and Cordero[l2] modified the above equation as follows

The coefficient c,, was evaluated numerically for ten strong ground motion records. On the basis of their study, an expression of the type given below was suggested [ 12]


in which is the average damping value. Figure 3.2 shows a comparison of the value of E,, obtained from Equations 3.18 and 3.19, with the average of numerically obtained values from ten earthquake records[ 121. Since Equation 3.19 is based on the average of c,,values obtained from several records, it is more appropriate to use it for a broad band earthquake input.


fi=O.l H Z , ~ = $ = O . O I .c 9)


- Formula Numerical 0

Frequency Ratio fj /fi Fig. 3.2 Comparison of the modal correlation coefficients from the formula with the average of numerical values from ten earthquake records[12].

Using the assumption of stationarity, Singh and Chu [13] derived an equation similar to Equation 3.13 from which an expression for E,. can be derived. Assuming earthquake motion to be white noise, Der Kiureghian[l4] also obtained an expression for E~ which is given below:

The double sum equation in which Der Kiureghian's expression is used has been called the complete quadratic combination (CQC)[15]. When the two modes have equal damping values, it can be shown that the E , values obtained from the Singh-Chu equation[l3] and those from Equation 3.20 (Der Kiureghian[l4]) are about the same. Both can be significantly lower than those given by Equation 3.15 (Rosenblueth and Elorduy[9]) and 3.18 (Gupta and Cordero[l2]) within the frequency ranges of interest. As will be shown later, these differences in the 8, values result in response variations that are not negligible. There is another important element in the expressions of the correlation coefficient which has not been explicity recognized in the published studies so far. Equations 3.15 and 3.18 are likely to overestimate the values of E,, when the damping ratios of two modes are sufficiently different. Consider a situation when w, and wJare large enough that the effect of the finite duration on the values of E,, in these equations can be neglected. Equations 3.15 and 3.18 can be approximated as follows:

An approximate form of Der Kiureghian's[l4] Equation 3.20 is

Equation 3.22 includes a coefficient, J(c,l;i)/c,,,that Equation 3.21 does not. The variation in the value of the coefficient with c,/c,ratio is tabulated below.


The coefficient is approximately unity when l,, and 6, are not very different. On the other hand it is much less than unity when 4, and 6, are sufficiently apart. This would have serious influence on the correlation coefficients for modes with close frequencies. Consider an example, w, = o,and l,,/l,, = 0.2. Equation 3.21 would give E,, = 1.0, and Equation 3.22, E, = 0.745. Further, consider the response of a secondary oscillator in resonant modes. Assume, R, = 1.01, R, = -0.99. The first value of E,,, Equation 3.21, would give R = 0.02, the second value, Equation 3.22 would yield R = 0.7 14. Our recent numerical experimentation using the actual earthquake motion data at North Carolina State University indicates that the correlation coefficient values and the resulting combined response values are relatively more realistic when the coefficient ,/(