PROCEEDINGS OF THE 3RD INTERNATIONAL WORKSHOP ON STRUCTURAL CONTROL
STRUCTURAL CUNTROL FDR CIVIL AND INFRASTRUCTURE
World Scientific
CIVIL AND INFRASTRUCTURE ENGINEERING
PROCEEDINGS OF THE 3RD INTERNATIONAL WORKSHOP ON STRUCTURAL CONTROL
STRUCTURAL CUNTRUL FUR CIVIL AND INFRASTRUCTURE ENGINEERING Paris, France 6-8 July 2000
Editors
Fabio Casciati University of Pavia, Italy
Georges Magonette Joint Research Centre Ispra, European Union
V f e World Scientific wfc
Singapore • New Jersey London • Hong Kong
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STRUCTURAL CONTROL FOR CIVIL AND INFRASTRUCTURE ENGINEERING Copyright © 2001 by World Scientific Publishing Co. Pte. Ltd. All rights reserved. This book, or parts thereof, may not be reproduced in any form or by any means, electronic or mechanical, including photocopying, recording or any information storage and retrieval system now known or to be invented, without written permission from the Publisher.
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Professor Grandori (left) receives his award
Professor Kobori (right) receives his award
Professor Masri (right) receives the award on behalf of Professor Housner
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Professor Kobori (left) with his wife and Professor Grandori (right) with his wife are congratulated at the end of the dinner in their honour
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Six photos taken during the dinner in honour of outstanding scientists in the area of technology innovation
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Six photos taken during the dinner in honour of outstanding scientists in the area of technology innovation
Institution The 3rd International Workshop on Structural Control was held in the new building of the Ecole Nationale des Ponts et Chaussees the oldest French School of Civil Engineering, located in Champs-sur-Marne. Steering Organization International Association of Structural Control (IASC) Steering Committee F. Casciati J-C.Chen S.F. Masri A.Nishitani
University of Pavia, Italy Hong Kong University of Science & Technology University of Southern California, USA Waseda University, Japan
Organizing Committee G. Magonette M. Collet F. Bourquin F. Marazzi F. Gualdi M.Valenti
ELSA Lab, ISIS, Joint Research Centre (JRC), Ispra, Italy LMARC, Besancon, France LCPC/CNRS, Champs-sur-Marne, France ELSA Lab, ISIS, Joint Research Centre (JRC), Ispra, Italy LMARC, Besancon, France LCPC/CNRS, Champs-sur-Marne, France
Sponsors European Association for the Control of Structure (EACS) Ecole Nationale des Ponts et Chaussees University of Pavia - Dipartimento di Meccanica Strutturale University of Genoa - Dipartimento di Ingegneria Strutturale e Geotecnica University of Naples - Dipartimento di Scienza delle Costruzioni Advantech Italy AD-Tech Italy JRC- ELSA Lab, Ispra
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Purpose The purpose of the periodic IASC (International Association of Structural Control) workshops is to provide an in depth update of the progresses accomplished in various regions of the word, in terms of theoretical developments, experimental validations and engineering implementations. This is achieved through the presentation of synthesis lectures and the participation in workgroups discussions. The lectures are open to invited participants selected to provide both a geographical cover and a broad technical spectrum. Active, passive, semi-active or hybrid control as well as non-destructive testing for health monitoring or damage detection will be addressed as well as related questions. The first act of the Association was to organize a restricted meeting, that was held in Honolulu, Hawai, in 1993. In the 1993 Workshop, 9 countries were represented with a total of 55 contributed papers. One year later, the 1st World Conference on Structural Control was held in Los Angeles. Four years after the Honolulu workshop, the 2nd International Workshop on Structural Control added the sub-title Next Generation of Intelligent Structures and collected 53 papers. The participants gathered in Hong Kong in December 1996, representing 11 countries. Two years later, in July 1998, the 2nd World Conference on Structural Control was held in Kyoto. On that occasion, the 3rd International Workshop was announced. In particular, three aspects were decided: the IASC Steering Committee, the workshop location and the Chairman of the Organizing Committee. The workshop was planned to pay special attention to the design of civil engineering structures by performance, to the structural control under strong earthquake and to the validation of the quality of experimental methods. The workshop gave all participants the possibility to establish contacts for further exchange of information, for promoting new initiatives and for improving future cooperations. It also provided the basis for the preparation of the Third World Conference on Structural Control, to be held in Como in April 2002.
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The editors have just three items to add: 1) the growth of interest on the topic of the workshop is evident when comparing the 9 countries represented in Honolulu and the 11 in Hong Kong with the 20 countries contributing these proceedings; 2) the papers were generally send in a form suitable for publication and thanks are due to all the authors for their kindness and accuracy. When necessary, a sort of maquillage was introduced by D. Miozzari, who spent four working weeks just to take care of these details; 3) the workshop was smooth for organisers and participants thanks to the impressive cooperations of three persons who at least require a citation: F. Marazzi, F. Gualdi and M. Valenti.
F. Casciati
Pavia, 20 October 2000
G. Magonette
CONTENTS Hosting Institution, Steering Organization, Steering Committee, Organizing Committee, Sponsors
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Purpose
xi
Fuoquet's Barriere Banquet in Honour of Outstanding Scientists in the Area of Technology Innovation G. Grandori G. W. Housner T. Kobori Working Group Reports: A) Control of Buildings and Bridges B) Modeling, Monitoring and System Identification C) Code Design Versus Performance Design An Identification Algorithm for Feedback Active Control N. D. Anh Optimum Seismic Response Control of Structural System Using Maxwell-Type Nonlinear Viscous Damper Based on Random Earthquake Response K. Asano, H. Nakagawa
3 7 9 15 19 25 27
39
Structural Control: Point of View of a Civil Engineering Company in the Field of Cable-Supported Structures M. Auperin, C. Dumoulin
49
Application of Control Techniques to Masonry and Monumental Constructions A. Baratta, O. Corbi, C. Cennamo
59
Modeling and Numerical Issues in the Active Control of Flexible Structures F. Bourquin, M. Collet, L. Ratier
73
Structural Control of Bridges: Recent Experiences in Cable-Stayed Bridges J. R. Casas
85
XIV
Some Implementation Issues in Structural Control S. Y. Chu, T. T. Soong, A. M. Reinhorn
93
Monitoring of Infrastructures in the Marine Environment A. Del Grosso
107
Seismic Retrofit of Church Buildings through Base Isolation A. De Luca, E. Mele, C. Giagnuolo
119
Towards a Safety Concept for Buildings with Structural Control U. E. Dorka
127
Semi-Active Control of 3-D Linear and Hysteretic Structures for Seismic Applications S. El-Borgi, C. Zammali, P. Tsopelas Report on 1999 Kocaeli and Diizce (Turkey) Earthquakes M. Erdik Health Monitoring and Optimum Maintenance Programs for Structures in Seismic Zones L. Esteva, E. Heredia-Zavoni Fuzzy Chip Controller Implementation L. Faravelli, R. Rossi
141
149
187
201
Health Monitoring of Civil Structures Using Spatial Information Obtained from Ambient Vibration Y. Fujino, M. Abe
215
Adaptive Control of MDOF Structures through Non-collocated Sensor/Actuator Pair V. Gattulli, F. Romeo
245
Designing and Testing Devices for Semi-Active Structural Control H. Gavin, M. Dobossy, J. Lamberton Cost Action F3 "Structural Dynamics" 1997-2001: Presentation and Some Preliminary Results J.-C. Golinval, P. Argoul
255
263
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Elastic Shear Frames with Tuned Liquid Column Dampers: A Controlled Study of Small-Scale Models A. Hruska, A. Dorfmann
281
Development of a Testing System for Dynamic Characterization of Energy Dissipating Devices H. Iemura, A. Igarashi, H. Tanaka
293
Performance of Variable-Damping Systems: Theoretical Analysis and Simulation J. A. Inaudi
301
A Benchmark Problem for Structural Health Monitoring and Damage Detection E. A. Johnson, H. F. Lam, L. S. Katafygiotis, J. L. Beck
317
Community Research Activities on Constructions: Competitive and Sustainable Growth RTD Programme G. Katalagarianakis
325
Outline of Safely Evaluation of Structural Response-Control Buildings and Smart Structural Systems as Future Trends K. Yoshikazu, T. Hiroyuki The Most Recent Applications of Seismic Isolation and Passive Energy Dissipation A. Martelli, M. Forni The State of the Art in Structural Control in Armenia and Proposal on Application of the Dynamic Dampers for Seismically Isolated Buildings M. G. Melkumyan
333
351
365
Earthquake Protection of Bridges Using Sliding Isolation System and MR Dampers S. Nagarajaiah, S. Sahasrabudhe, R. Iyer
375
Recent Developments in Smart Structures Research in India S. Narayanan, V. Balamurugan
385
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Semiactive Structural Control Strategy for Buildings against Severe Earthquakes A. Nishitani
407
Preservation of Mediterranean Historical Monuments Using Innovative Seismic Techniques A. Osman, A. Saleh, A. El-Attar
417
Simple Controller Design for Nonlinear Structural System Using Robust Performance Property W. Park, H.-M. Koh, D.-H. Ha
437
Perspective of Application of Active Damping of Cable Structures A. Preumont, F. Bossens
445
Structural Control of a Rotating Spacecraft with Elastic Spike Antennas Using the Magnetohydrodynamic Control System B. Rabinovich Activities of the European Laboratory for Structural Assessment in the Field of Structural Control for Civil Buildings, Bridges and Architectural Heritage V. Renda, G. Magonette, J. Molina, D. Tirelli, F Marazzi
453
463
Optimal Design of Dampers and Stiffeners in Structures Using Active Control Approaches A. M. Reinhorn, N. Gluck, J. Gluck, R. Levy
473
On a Stability-Based Approach for Robust Active, Semiactive and Hybrid Structural Control J. Rodellar
487
Civil Structures in the New Millennium Mixing Art, Nature, and System Science R. E. Skelton
495
Parametric and Nonparametric Adaptive Identification of Nonlinear Structural Systems A. W. Smyth, S. F. Masri, E. Kosmatopoulos, A. Chassiakos
521
The Athens (Greece) Earthquake of September 7, 1999 C. A. Syrmakezis, A. A. Sophocleous
529
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Controlling Modeling Error Impact in Structural Parameter Estimation S. Wadia-Fascetti, S. Ozgu, M. Sanayei Active Control Requirements in Railway Projects H. Wenzel
543 551
Semi-Active Base Isolation Control of a Building Using Variable Oil Damper K. Yoshida, T. Fujio
567
Participant Distribution List
577
Author Index
Fouquet's Barriere Banquet.
In honour of outstanding scientists in the area of technology innovation:
GIUSEPPE GRANDORI GEORGE W. HOUSNER TAKUJI KOBORI
Address: Fouquet's Barriere 99, avenue des Champs Elysees 75008 PARIS- France Tel. (33) 01 47 23 50 00
Le Founquet's Barriere in Paris On 8th November 1899, Louis Founquet opened at number 99, avenue des Champes Elysees a restaurant which was to become for the entire world the restaurant of the Paris elite and required meeting place for the greatest international stars.
PRESENTATION FOR PROFESSOR GIUSEPPE GRANDORI REMARKS BY FABIO CASCIATI Department of Structural Mechanics University ofPavia E-mail:fabio @dipmec. unipv. it Professor GIUSEPPE GRANDORI, Pippo for the friends, is presently a retired person enjoying life. But his ability to concentrate himself on a specific problem, to identify the simplest solution and to lead the team in charge of implementing it, remain in the mind of a generation of Italian academic colleagues, some of them having Prof. Grandori as direct advisor, some others, as me, having his as indirect advisor. Indeed he "invented" Earthquake Engineering in Italy and was able to convince the government authority that it exists. Of course for this he achieved numerous accomplishes. But to celebrate his career this evening is out of the scope of this meeting. Indeed we are running the 3rd International Workshop on Structural Control; we have just survived the 2nd European Conference on Structural Control and we are looking for a 3 rd World Conference on Structural Control in Como, Italy, two years from now. But, which was the seed of all this activity? No mater than the adoption in civil engineering of schemes providing an intelligent counteraction to external actions is due to Kobori in Japan. No matter that Housner in the USA was attracted by the idea and promoted its extension to any form of intelligent behaviour. But this is the first workshop in Europe. So let me try to go back: I see a lot of passive control implementation in building and bridges. From where are they coming? European research programs are now emphasising the development and demonstration technologies addressing "rehabilitation, upgrading and strengthening of civil and industrial structures and facilities". I went back to 1976 to find something that resembles a seed: the four drawing which complete this short report. Prof. Grandori initial studies had been focused on plasticity theory. They date 1952, 1953 and so on, but in 1976 he had the idea to exploit this special behaviour. In a personal communication he wrote me in Italian: "... I was working on a constraint device for long monolithic bridge decks in seismic areas. It is a constraint providing the horizontal reaction in the longitudinal direction (as long as the other supports do not work on the longitudinal degree of freedom). During a strong earthquake, for a deck of 100 meters and more, (there were bridges of length 900 m built in Friuli with this technology) the resulting axial force of this single side constraint could reach significant values, oscillating from tension to compression. In a pre-stressed concrete deck this can result in a saddle failure". 3
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Figure 1 - Longitudinal view of the constraint device by Professor Grandori.
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Figure 2 - Way of working of the constraint device by Professor Grandori.
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Italy had just suffered the Friuli earthquake and the motorway network company (Societa Autostrade) realised the problem. The one in Figure 1 is the solution given by Prof. Grandori: it works on the axial force value as summarised in Figure 2.
Figure 3 -Hysteretic cycle for a couple of Nicuage steel elements. As it occurs for many innovative ideas, Professor Grandori, starting from the initial goal of limiting the axial force, realised that he was introducing a passive damper in the design of Civil Infrastructure Systems. A single couple of Q shaped Nicuage steel elements can easily be tested in the laboratory and provides the hysteretic diagram of Figure 3. When three couples are mounted in series as proposed by Professor Grandori, the result of Figure 4 is achieved and one realises that the device also works as a damper.
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Figure 4 - Hysteretic cycle of the constraint device proposed by Professor Grandori. For his pioneering efforts toward controlled structures in Europe, IASC would like to honour him tonight. References 1.
Grandori G. (1952), Contributo alio studio deU'equilibrio elastoplastico (in Italian), Proceedings of the Milan Engineer Society, 1-2, 3-11.
2.
Grandori G. (1954), Risultati di un'esperienza riguardante il comportamento delle strutture elasto-plastiche (in Italian), Proceedings of the Milan Engineer Society, 3-4, 1-3.
3.
Grandori G. (1956), Un'esperienza sul comportamento di travi inflesse elastoplastiche (in Italian), Proceedings of the Symposium on Plasticity in Structural Engineering, Varenna, Italy, 1-7.
PRESENTATION FOR PROFESSOR GEORGE W. HOUSNER
REMARKS BYSAMIFMASRI Department of Civil Engineering University of Southern California E-mail:Mastri @ use. edu
Good evening. As we attend this banquet to honor some of our renowned scientists for their outstanding contributions to Structural Control, I have been asked to say a few words of introduction for Professor George W Housner, CF Braun Professor Emeritus of Engineering at the California Institute of Technology. Professor Housner was a native of Michigan and received his undergraduate education at the University of Michigan; and he received the Ph.D. degree from California Institute of Technology. He was one of the pioneers in earthquake engineering and he has served as President of the International Association for Earthquake Engineering, the Earthquake Engineering Research Institute, and the Seismological Society of America. Professor Housner has done research in earthquake engineering on the characteristics of destructive ground shaking, earthquake design spectrum, seismic design of structures including high-rise buildings and liquid storage tanks. He served as Chairman of the California Governor's Board of Inquiry on the Loma Prieta Earthquake; Chairman of the California Seismic Advisory Board of Caltrans and Chairman of the Earthquake Analysis Board of the California Department of Safety of Dams. He has served as consultant on design of the Trans-Arabian pipeline, the California Water Project, the San Francisco Bay Area Rapid Transit system, etc. Professor Housner's contributions to earthquake engineering have been recognized by election to membership in the US National Academy of Sciences and the National Academy of Engineering, the Japan Academy of Science, and the Indian Science Academy. He has received many honors and awards, including the National Medal of Science (the highest civilian award in the USA), the Von Karman Medal of the American Society of Civil Engineers, and the Society Medal of the Seismological Society of America. One would think that with all of the awards and accolades that he has received, and the numerous accomplishments that he has achieved, he would have retired a long
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time ago, and essentially gone fishing or done something equivalent. However, as many of you know, he continues to go to his office on a daily basis, and his work habits would put many assistant professors to shame! When we approached him about 10 years ago (when he was at the young age of 80 years), to help us lead the US national effort in the area of Structural Control, he did not hesitate to accept this leadership role once he understood the broad ramifications of the field and its great potential to lead to safer and more efficient structures,. Professor George W. Housner early perceived the potential benefit to earthquake engineering from research on structural control and decided that there should be an International Association for Structural Control (IASC) and was its first president. For his great and pioneering efforts in the birth and nurturing of the general field of Structural Control in the USA as well as through out the world, IASC would like to honor him tonight.
PRESENTATION FOR PROFESSOR TAKUJI KOBORI REMARKS BY AKIRA NISHITANI
Department of Architecture, Waseda University, Tokyo E-mail:
[email protected] Dr. Takuji Kobori, Professor Emeritus of Kyoto University, is one of the most distinguished researchers in structural engineering and earthquake engineering. He has been leading the international research community for more than half a century. The International Association for Structural Control was officially established in 1994 with the great leadership of Professor Kobori and Professor George W. Housner, the founding president of IASC. Professor Kobori succeeded Professor Housner as President in 1996, serving as President for four years until July of 2000. It gives me a great pleasure and honor to introduce such a distinguished researcher, Professor Takuji Kobori, on this specific occasion. Professor Kobori was born in Tokyo in 1920. He graduated from Waseda University in Tokyo, Japan, with a B.S. degree in architecture. As it may be widely recognized, the Japanese way of architectural education provides students with not only architectural design and planning courses but also structural engineering courses. Finishing his graduate study at Waseda, he was appointed as a faculty member of Architectural Engineering at Kyoto University in 1951. Since then he has substantially and deeply influenced the theoretical and practical developments of structural engineering publishing more than 350 papers on structural dynamics and vibrations. Shortly after he became Professor Emeritus of Kyoto University resigning from the teaching there, he was invited to join Kajima Corporation as Executive Vice-President in 1985. Then he worked as Executive Vice-President until 1992 and he is now Chief Executive Advisor of Kajima Corporation. Moreover, since he founded Kobori Research Complex in 1986, he has been the president. Professor Kobori has made a lot of contributions to the activity of the International Association for Earthquake Engineering (IAEE) for a long time. He was a member of Board of Directors for this association from 1988 to 1996 and he is now an Honorary Member of IAEE. He is also one of the leaders in the entire field for architecture in Japan. From 1983 to 1984 he took the leadership as President of the Architectural Institute of Japan (AIJ), which is the national association covering the whole areas for architecture such as architectural design and history, urban planning, and structural, environmental and construction engineering, etc. He is now Honorary Member of AIJ. 9
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Professor Kobori has dealt with a wide range of problems in the fields of structural engineering and earthquake engineering in his long academic career. With full recognition of his outstanding contributions to the advancement of earthquake engineering, the Architectural Institute of Japan awarded him the AIJ Grand Prize, the highest and most honorable award of the Institute, in 1990. His achievements are really enormous, covering a variety of issues on structural dynamics and vibrations. It would be truly impossible for me to describe everything. I am afraid I could present only a few of his prominent achievements. Professor Kobori is one of the pioneers in nonlinear vibrations for structures under seismic excitation. Giving a thorough treatment to the analytical investigations of structural nonlinear vibrations, he [1,2,3,4] clarified the significance of the nonlinearization of structural behavior in seismic event and indicated that a combination of vibrational energy dissipation and elongation of the natural period resulting from either its inherent or artificially-added nonlinearization brought about the seismic response reduction. His creative research achievements in this field have established the foundation for aseismic design of high-rise buildings in Japan. Until the recent enforcement of the new Building Standard Law of Japan, all of the highrise buildings in Japan had to be assessed by the High-Rise Building Appraisal Committee at the Building Center of Japan. This committee had been chaired by him for long time. In addition, Professor Kobori is one of the fathers of structural control. As early as in 1956, he proposed the necessity and significance of incorporating the concept of automatic control in the earthquake resistant design of building structures [5]. This idea was extended in his subsequent papers published in 1960 [6,7]. In these papers the fundamental principles for structural control were presented together with the Japanese expression equivalent to "Seismic-response-controlled structures" in English. Since that time he has been instrumental in the theoretical development and practical implementations of structural control. Finally, the world's first activelycontrolled building [8,9] was designed by him in Tokyo, Japan, in 1989. The birth of this building really appealed the attentions of many research and practicing engineers worldwide. In a way, the birth of this landmark building, I would say, led to the establishment of IASC in 1994. What I presented above is only a few of Professor Kobori's achievements. As I mentioned already, his academic contributions to structural engineering are really enormous and outstanding. I cannot therefore present everything. But I can say one thing. Without his seminal contributions the field of structural engineering would be entirely different today.
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References 1. Kobori, T. and Kaneta, K., Introduction to nonlinear vibrations of seismicallyexcited structures, Trans. Architectural Institute of Japan 51 (1955) pp.50 - 60 (in Japanese). 2. Kobori, T. and Minai, R., Nonlinear structural vibration subjected to the earthquake loading (Part 1 Natural nonlinear response process), Trans. Architectural Institute of Japan 51 (1955) pp.61-69 (in Japanese). 3. Kobori, T. and Minai, R., Nonlinear structural vibration subjected to the earthquake loading (Part 2 Artificial nonlinear response process), Trans. Architectural Institute of Japan 52 (1956) pp.41-48 (in Japanese). 4. Kobori, T., Nonlinear structural vibration subjected to the earthquake loading (Part 3 Control, filter and earthquake loading pattern recognition), Trans. Architectural Institute of Japan 54 (1956) pp.401-404 (in Japanese). 5. Kobori, T., Quake resistant and nonlinear problems of the structural vibrations to violent earthquake, J. Kyoto University Disaster Prevention Laboratory; 5th Anniversary Edition (1956) pp. 116-124 (in Japanese). 6. Kobori, T. and Minai, R., Analytical study on active seismic response control: Seismic-response-controlled structure 1., Trans. Architectural Institute of Japan 66 (1960) pp.257-260 (in Japanese). 7. Kobori, T. and Minai, R., Condition for active seismic response control: Seismic-response-controlled structure 2., Trans. Architectural Institute of Japan 66 (1960) pp.253-256 (in Japanese). 8. Kobori, T. et al., Seismic-response-controlled structure with active mass driver system (Part 1) Design, Earthquake Engineering and Structural Dynamics 20 (1991)pp.l33-149. 9. Kobori, T. et al., Seismic-response-controlled structure with active mass driver system (Part 2) Verification, Earthquake Engineering and Structural Dynamics 20(1991)pp.l51-166.
Working Group Reports
WORKING GROUP REPORT: CONTROL OF BUILDINGS AND BRIDGES WG CO-CHAIRPERSONS: L.A. BERGMAN University of Illinois at Urbana-Champaign J.R. CASAS Universitat Politecnica de Catalunya K. YOSHIDA Keio University WG RECORDER: S.J. DYKE Washington University in St. Louis
Summary of Previous Benchmark Activities I s ' Generation •
3-story test buildings based on linear experimental building (Published in a special issue of Earthquake Engineering and Structural Dynamics, September 1998, http://www.nd.edu/~quake/).
2nd Generation •
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20-story full scale linear, seismically excited building (published in the Proceedings of the Second World Conference on Structural Control) available at: http://www.nd.edu/~quake/ 76-story full scale linear, wind excited building (published in the Proceedings of the Second World Conference on Structural Control) available at: http://www-ce.engr.ccny.cuny.edu/people/facultv/agrawal/benchmark.html
3rd Generation • •
•
3-, 9-, and 20-story nonlinear, seismically excited building, available at: http://www.nd.edu/~quake/ 76 story full scale linear, wind excited structure with improved wind models, available at: http://www-ce.engr.ccny.cunv.edu/people/facultv/agrawal/benchmark.html Full scale cable-stayed bridge (recently completed), available at: http://wusceel.cive.wustl.edu/quake/ 15
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•
Full scale base isolated linear building (in progress)
Motivation Behind Benchmark Problem Development • •
To have a common basis for comparing methodologies To provide realistic numerical testbed for examining issues
Advantages Offered by Benchmark Problems • • •
Pedagogical aspects - training of students Educating practitioners in state-of-the-art in structural control Potential for collaboration
Main Objectives/Goals 1.) Questions for Discussion The working group was asked to consider the following questions for discussion. • Should there be a fourth generation of benchmark problems and what should they be? • Should we recommend continuation of the incremental approach, or is there a particular class of problems that needs to be exhaustively studied? • Are we ready for a simultaneous structure/control design exercise? 2.) Issue to be Considered in Future Benchmarks The following issues were identified as important issues to include in the development of future benchmark problems in structural control: • Amplitude dependence (for nonlinear structures and/or control systems) • Robust performance and stability measures (considering structural degradation, modeling errors, environmental factors, etc.) • Near-field earthquakes (e.g., Turkey, Taiwan)
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Recommendations of Working Group 1.) Regarding web site development: • Recommend development of a 3rd generation benchmark results web site on the wind excited building problem • Recommend development of a 2nd generation benchmark results web site on the seismically excited building problem 2.) Add robustness measures for authors to test robust stability and performance of their designs 3.) Add comments to existing benchmark to: • Encourage authors to consider amplitude dependence in weighting evaluation criteria • Encourage authors to compare results to comparable passive systems to clearly demonstrate any advantages of active or semi-active systems 4.) Possible Ideas on New Benchmark Problems • 3D base isolated building with a nonlinear superstructure and a choice of isolation bearings with a focus on near-field earthquakes • Possibly an integrated structure/controller benchmark in the future as a separate problem.
List of Partcipants in the Benchmark Control Problems Working Group Meeting NAME Lawrence A. BERGMAN Franco BONTEMPI Joan Ramon CAS AS Shirley DYKE Kazuo YOSHIDA Narito KURATA Jose INAUDI Luciana BARROSO Hirokazu IEMURA Henri GAVIN Jose RODELLAR Andre PREUMONT
E-MAIL lber gman @ ui uc. edu
[email protected] [email protected] [email protected] [email protected] [email protected] archis ©onenet.com.ar
[email protected] [email protected] [email protected] [email protected] apreumont @ ulb. ac. be
WORKING GROUP REPORT: MODELING, MONITORING AND SYSTEM IDENTIFICATION WG CO-CHAIRPERSONS: J. BECK California Institute of Technology Y. FUJINO University of Tokyo G.MAGONETTE Joint Research Centre of European Commission WG RECORDER: ALESSANDRO DESTEFANO Politecnico di Torino
The working group collected a wide range of experiences and cultural backgrounds. Therefore it is a difficult, but stimulating task, to define and fix common goals. In the same time, that variety of experiences and cultural backgrounds offers a good occasion to compare different conceptual worlds and interests for people having, anyway, at list a common language.
Main Objectives and Goals
Inside the whole activity of the Working Group it is possible to recognize three main general goals and three corresponding research topics. They are listed above in a synthetic way:
1) Health monitoring as a complex design of strategies to keep under control given families of structures in condition of possible strong loading and evolutive health state. A first interesting contribution has been given during the workshop by Del Grosso, about infrastructure in marine environment. Here the possibility of replacing with the environmental event statistics the direct, more difficult, structural health monitoring is stressed out, leading to a drastic
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simplification of control strategies. A second contribution, by Esteva, shows the needing of performing a cumulate damage model assessment to obtain a residual safety and reliability monitoring for structures subject to strong earthquakes. 2)
Damage
detection
and
localization
through
dynamic
testing
3) General mechanical characterization of structure to assess in a precise and reliable way structural models to be used for the design and exploitation of structures and control systems. Goals 2 and 3 involve some common methodologies and problems
Modeling and Testing
The research topics related to goals 2 and 3 are discussed together, since they present several common aspects. In both domains a mathematical model is generally required. A model can be linear, or locally linearized, or non-linear. In the first two cases the model is parametric, often modal, and modal parameters (frequencies and shapes) are updating and assessment tools.Is modal damping suitable? No unanimous answers exist up to now . Model uncertainties have to be kept into account to avoid identification errors. Local linearized mode shapes can be excitation amplitude dependent. Modal parameters can be strongly influenced by temperature changes (see the contribution by Fujino, Aktan and the results of the COST Project). Modal testing for damage location requires higher mode parameters assessment and dense and consistent set of measure points. Experimental modal identification is a fundamental tool for model correction. Time domain and time frequency domain are modern approach tools mostly for ambient vibration. In the third case, if non-linearities are moderate, both parametric and non parametric approaches can be used (Smyth).
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The complexity of non-parametric approaches (Volterra -Wiener series) makes it useful to look for the help of heuristic tools as neural network for an operative use. Reverse path method and proper reduction strategies have been tested in the COST project. A nonlinear model assessment has been applied to simple structures or to more complex ones but only with localized non-linearities Strong non-linear hysteretic behavior is associated with cumulate damage and damage index evolution (Esteva), and memory supplied models are a need.
Modeling Problems
Model updating and assessment require strategies to keep into account the geometric model errors and to separate them from other causes of errors and uncertainties. (Wadia Fascetti) Special purpose studies related to specific modeling problems are shown in the contributions by Anh and Renda. The first one is devoted to the feedback model assessment for active control and the related problem of time delay. The second-one points-out the correction to be given to non linear special materials characterization, strain rate sensitive as rubber bearings or shape memory alloys, when non real time experimental tests, as pseudo-dynamics does, are performed on a structure.
Testing Conditions
Testing conditions can be generally grouped into two main classes: 1) Ambient vibration: often a response analysis is performed using outputonly experimental data. Such testing conditions class is characterized by low cost and low impact on the service life of the structure. Its
22
effectiveness for health monitoring purpose is still an open problem. It leads generally to parametric (modal) identification with local linearization. It has been shown how parameters can be sensitive to response amplitude and environmental conditions (Fujino, Aktan). 2) Strong excitation: This class includes artificially excited tests and strong earthquake or wind records. A higher level of energy and response amplitude requires one to handle a non-linear reference model. They can lead to a direct assessment of the residual strength and structural safety (Esteva).
Benchmarks Benchmarks are fundamental tools to verify the effectiveness of different methods and approaches. In a working Group having a so wide spectrum of methodologies and goals, they can offer a unifying desk and common objects to allow the participant to exchange experiences. They represent very important contributions to reach general goals 2 and 3. Benchmark case studies have been set up in the US and already used as comparative testing base (Beck). Four steel braced frames were built and tested, with localized defects in different condition of excitation and measure. The results and comparative comments have been shown (Beck). A free access to the databases has been offered to all researchers belonging to the Working Group. Benchmarks data from European COST Project (Golinval) have been shown and the results compared and commented (Argoul). Data -bases, even if not yet declared as openly accessible to everyone, should be obtainable on request from the author. Two databases of experimental data on existing bridges have also been offered on request, without any restrictive condition (Aktan and Wenzel).
23 Recommendations of the Working Group The panel discussion of the working group was concluded by the following list of recommendations to the WG members themselves and to the IASC board: 1) Full future accessibility is required to benchmark and experimental data from all over the world. 2) Separate databases and benchmarks should be prepared for buildings and bridges. 3) Separate sets of benchmark and testing data should be made available for ambient and strong excitation data. 4) If available, data (Actually it seems that no complete and reliable data exist under very strong earthquakes) on instrumented buildings under strong earthquakes should be made accessible 5) A web page should be created and kept update by IASC to contain links to all available data set sites or reference persons. 6) A special session of the next World Conference on Structural Control should be devoted to the activity of the Working Group. List of Participants in the Working Group Meeting NAME
E-MAIL
Sara WADIAFASCETTI Jim BECK
[email protected] Vincenzo GATTULLI Raimondo BETTI
Gattulli @ ing.univaq.it
Scott JOHNSON
Smj 1 @cec. wustl.edu
Juan Martin CAICEDO Alessandro DE STEFANO Pierre ARGOUL
Jcl 1 @cive. wustl.edu
[email protected] Betti @ civil.columbia.edu
[email protected] [email protected] Andrew SMYTH
[email protected] Luis ESTEVA
Este va @ merlin.iingen .unam.m X
Georges MAGONETTE Yozo FUJINO
[email protected] Erik A. JOHNSON Helmut WENZEL
[email protected] Luis DORFMANN
[email protected] [email protected] [email protected] WORKING GROUP REPORT: CODE DESIGN VERSUS PERFORMANCE DESIGN WG CO-CHAIRPERSONS: A. DEL GROSSO University of Genoa Y. KITAGAWA Keio University T.T. SOONG State University of New York at Buffalo WG RECORDER: U. DORKA University of Rostock
Main Objectives and Goals 1.) Code developments related to structural control Present, code developments concern base isolation (that is nearly to be included as a code-based design for retrofitting of existing structures and designing of new structures in most countries), and energy dissipation devices (that are already included as standard practice in some code and are the subject of experimental regulations in other countries). No development is foreseen in the near future for semi-active and active devices. The WG has recognized that the application of control techniques has been made in Japan and other Asian Countries outside codes, on a peer review basis, and, in the opinion of the WG, IASC should try to first encourage the development of real projects also in other countries, forming review panels to promote applications outside codes and creating a reference for building authorities. 2.) Topics of specific interest for application Considering that experimentation should be the basis for encouraging applications and code developments, the WG has recognized that TMD and AMD have already reached a critical mass in real constructions. Their use is therefore ready to be spread worldwide. The WG encourages the collection and diffusion of detailed information concerning design, peer review, construction and operation of structures equipped with such devices, already built in Asia. 25
26
3.) Performance design The WG has recognized that performance-based design, in terms of reaching a target reliability for service and ultimate limit states, is still far away to be introduced as a standard practice in civil structural engineering. However, from the point of view of the development of such a practice in the field of control, the problems related to performance of devices shall be considered before taking into account the performance of complete structural systems. The WG recognizes that applications of energy absorbing systems are already spread enough to validate theoretical reliability models.
Recommendations of the Working Group Based on the previously stated findings, the WG indicates the following research lines as capable of influencing the development of code related activities: 1) Benchmarking studies; 2) Code-related research, including collection, analysis and dissemination of information concerning actual implementation of passive and active systems; 3) Performance evaluation of all the components that should be included in a controlled system; 4) Development of reliability models for components and controlled systems. Research on the above topics should therefore be encouraged by IASC.
List of Participants in Working Group Meeting NAME Veit BAYER Andrea DEL GROSSO Uwe E. DORKA Yoshikazu KITAGAWA Andrei REINHORN Andrew SMYTH Tsu T. SOONG
E-MAIL
[email protected] [email protected] [email protected] Kitagawa @ sd.keio.ac.jp
[email protected] [email protected] [email protected] AN IDENTIFICATION ALGORITHM FOR FEEDBACK ACTIVE CONTROL
N.D. ANH Institute of Mechanics 264 Doi Can, Hanoi, Vietnam E-mail:
[email protected] The aim of the paper is to present a control law for feedback active controlled structures in which a control algorithm is proposed to identify the external excitation with a time delay. The time duration in which the external excitation acts on the structure is devised into n subintervals. In each subinterval the external excitation can be identified and then is selected with opposite sign as the control force for the next subinterval. The realization of the identification control algorithm in the practical application mainly depends on the computing capacity of the involved computer and requires an investigation with respect to its robustness and stabilization.
1
Introduction
In recent years much progress and new concepts have been archived in the development of the structural control in reducing the response during the excessive vibrations due to environmental loading such as earthquakes and winds. Many control devices have been installed or used temporarily in full-scale structures to provide enhanced structural behavior for improved service and safety Casciati et al. [2], Housner et al. [6], Kobori [7]. The control of structural motions due to environmental excitation can be done by various means. Among them passive and active control systems represent fundamental approaches in the use of control strategies for response reduction in structures. The first approach proposes to reduce the energy of the environmental disturbances acting on the structure by transferring its part into other structural elements (secondary systems) while the latter uses the control counterforces produced by actuators to balance the environmental loading. Since in the first approach the structure does not receive any additional energy, the application of passive devises is usually effective up to a certain limit. A major disadvantage of the second approach is that in many cases of practical interest, the required additional forces are too large for existing actuators (for example, the case of earthquake excitation). Thus, the third approach which presents the combination use of passive and active control systems leads to so called hybrid control systems as an adequate means of structural protection with high level of performance against dynamic hazards. Since the objective is to control the structural response and external disturbances are a main resource producing it, one may expect that the control problem would be more simplified if both structural response variables and external
27
28
disturbances could be measured during the structural motion. Thus, the case where the information on both the structural response quantities and external excitations are used for designing the control forces, leads to so called feedback-feedforward (or closed-open loop) active control. In fact, as it has been justified by many researchers, the effect of feedback control is to modify the structural parameters so that it helps the structure to react to the external disturbances in a more favorable way while the feedforward components balance or totally eliminate the external loading. In the classical active control the active control forces can be used not only to reduce the structural response but they can be chosen in an optimal way. If the active controlled structure can be modeled as a linear system and the control force vector is selected as a linear combination of measured displacement, velocity and excitation vectors then the conventional optimal control law is that minimizes a performance index which is defined as the integration of a quadratic expression with respect to the state and control vectors Yao [18], Soong [14]. It seems that the active control problem for linear structures can be solved perfectly if there is not the following situation: it is usually that one is unable to measure the external excitation while the structural response can often be measured. Hence, it is obvious there is a need to look for new control laws to overcome this deficiency of the classical optimal active control algorithms as above described. Many new and interesting control strategies and concepts have been proposed such as reliability-based control Spencer et al. [16], bounded state control Lee and Kozin [9], Del Grosso and Zucchini [3], parametric and predictive control Lai and Wang [8], Rodellar et al. [13], adaptive control based on the intelligent devices using fuzzy logic and neural networks Casciati, Faravelli and Yao [1], Faravelli and Venini [4], Fu [5], Widrow and Lehr [17]. The aim of the paper is to present a control law for feedback active controlled structures. A control algorithm is proposed to identify the external excitation with a time delay.
2
Problem Formulation
Suppose that the feedback active controlled structure is modeled by an n-degree-of freedom system whose equation of motion is described by the nonlinear differential equation
Mz(t) + G(z(t),z(t)) = u(t) + f(t)
(i)
where M is the nxn mass matrix, z(t) = (z1(t),z2(t),...,zn(t)) is the nxl displacement vector, G is a nonlinear function of displacement and velocity components, f(t) = (f'(t),f2(t),...,fa(t)) is the nxl vector representing external excitation. Since the feedback active control is considered here, thus, the nxl control force vector provided by actuators u(t) is selected as a function of measured response. As the
29 structural response mainly depends on the structural parameters characterizing the internal dynamics of the structure and on the external disturbance, we consider first the structure free of external excitation and control force Mz(t) + G(z(t),z(t)) = 0
(2)
The free vibration of the system (2) is described as z(t) = z 0 ( t , z ( t 0 ) , z ( t 0 ) )
(3)
where z(t0), z(t0) are initial displacement and velocity vectors of the structure. Suppose that the structure is designed in such a way that one has
||z0(t,z(t0),z(t0)||
< 5X
(4)
if
|z(t 0 )|| and |z(t 0 )| < e
(5)
where || . || is a norm, 8i and e are small positive numbers. To archive (4) one may use methods of increasing the energy dissipation capacity to reduce the amplitudes of free vibration of the structure. At present, many passive energy dissipation devices such as friction sliding and metallic yield dampers have been installed in structures in order to enhance energy dissipation in structural systems Soong and Dargush [15]. In particular, one may use a part of the active control force in the form of linear velocity feedback u f ( z ( t ) ) = -Sz(t)
for increasing damping in the structure. In terms of the stability theory it is supposed that the equilibrium position of the system (2) z = 0, z = 0 is stable in Lyapunov sense. Under this assumption the control problem for the system (1) is to find the active control force u(t) necessary to keep the displacement norm below specified bounds in the presence of any disturbances less than a certain value provided that the initial values of the displacement and velocity norms are sufficiently small. In the mathematical description it means that find u(t) = u(t, z(t), z(t), z(t)) such that ||z(t)| < 8 if | z ( t 0 ) | , | z ( t 0 ) |
/7 {t) + knI(t) + cn(pI(t) + kn(p,(t)+mn(i>II{t) + cl2(pII{t) + ku(pI1(t) =
=
-r'xUgxii)-r'yugy{t)-f's
m21cp; (r) + c2l(p, (t) + k2l(p, (t) + m22ipn {t) + c22a{t)+Cein{t) + Ke J-
-
> ~
S
> S
>
Vd
S
where T stands for a 2nd order high-pass filter, with a 0.1 Hz cut-off frequency. The fundamental frequency of the beam is 2.5 Hz. Now, computing M^ from displacement measurements is not necessarilly easy but we may take advantage of the dynamical properties of the normal modes to define robust and easily implementable formulas for the right-hand side of the controller equation. This will be explained in a more detailed paper. A global similar efficiency as with the primal formulation is obtained but there is no more spill-over at "high frequencies". The efficiency still remains limited by the necessary high-pass filtering. Loosely speaking, this "dual" control law proves slightly less efficient for the fundamental mode than the "primal" one but behaves nicer at higher frequencies. 5
Mixed primal dual control synthesis
Based on above conclusions, the new idea writes as follows : by just using each feedback in the frequency range where it is efficient and where there is no spill-over effect, provided both regions overlap, and they do, one could certainly increase the gain and ask for more efficiency. We propose here the additive blending v
= 2 $PVP
+
GdVd)
where vp denotes the control obtained with the primal formulation and Vd denotes the control obtained from the dual one, and where Qp and Gd stand for gentle 1st order filters defined here as
^( S ) = TTT:
gd{s)=
lf^
»r = 10Hz
"d = 20Hz
81
OP
capteur 3
Filtre modal
^ derivation
Calcul des contributions du moment et de sa derive
capteur 2
Gain de Komomik primal
Gain de Komomik dual
capteur 1
(filtre passe bas )
(Double integration )
X
capteur Potvjbniil
F i g u r e 5. e x p e r i m e n t a l s e t t i n g for t h e p r i m a l d u a l s y n t h e s i s
. xio
TEMPS (s)
Figure 6. primal dual synthesis : the first 3 modal contributions to the response of the beam, u = 5
As anticipated, the parameter ui can now be increased and a much better efficiency is observed. In half a second all the vibrations of the beam get damped. The new higher efficiency limitation seems to be due to above choice
82
TEMPS (s)
Figure 7. primal dual synthesis : the "weak energy" of the controlled beam, ui = 5
of weighting parameters. We then expect to increase the potential efficiency by just modifying above decomposition. In terms of computational time, the resulting strategy is only twice as expensive as the previous one. 6
Concluding remarks • An apparently new way of blending mechanical formulations and associated low and high pass filters to define control laws proves much more efficient than previous strategies based on a single mechanical formulation. This remark may open the way to new classes of control algorithms that are both efficient and practically stable. • In particular, the mixed primal dual control synthesis may be of interest in the case of high modal densities which is the case of many, if not all, complex structures. • A thorough analytical investigation is needed to fully understand and enhance the properties of the new class of control algorithms.
83
• The dual formulation shares common features with IFF x . For example, stresses are "measured", and finally time-integrated. Therefore combining IFF and DVF might also be of interest. References 1. Y. Achkire and A. Preumont. Active tendon control of cable-stayed bridges. J. Earthquake Engineering and Structural Dynamics, 25:585597, 1996. 2. F. Bourquin. A numerical approach to the exact controllability of eulernavier-bernoulli beams. Proceedings of the First World Conference on Structural Control, Pasadena (California), pages 120-129, 1994. 3. F. Bourquin. Approximation for the fast stabilization of the wave equation from the boundary. In Proceedings of MMAR2000, Poland, 08-2000, 2000. 4. F. Bourquin, J-S. Briffaut, and M. Collet. On the feedback stabilization: Komornik's method. In Proceedings of the second international symposium on active control in mechanical engineering, Lyon, France, 1997. 5. F. Bourquin, J-S. Briffaut, and M. Collet. Smoothed fast stabilization. In Proceedings of the second world Conference on structural control, Kyoto, 1998. 6. F. Bourquin, J-S. Briffaut, M. Collet, M. Joly, and L. Ratier. Fast control algorithms for beams : experimental results. In Proceedings of Forum Acousticum 1999, Berlin, 1999. 7. F. Bourquin, M. Collet, M. Joly, F. Lene, and L. Ratier. An efficient control algorithm for beams : experimental results. In Proceedings of ACTIV1999, Florida, 12-99, 1999. 8. J-S. Briffaut. methodes numeriques pour le controle et stabilisation rapide des structures. PhD thesis, Ecole Nationale des Ponts et Chaussees, Paris, 1999. 9. V. Komornik. Rapid boundary stabilization of linear distributed systems. Siam J. Control and Optimization, 35(5):1591-1613, 1997. 10. G. Lebeau. Controle de l'equation de schrodinger. J. Math. Pures Appl., 71:267-291, 1992. 11. L. Ratier. stabilisation rapide des structures et raise en oeuvre experimentale. PhD thesis, ENS Cachan, France, 2000. en preparation. 12. J. Urquiza. Controle et stabilisation des structures haubanes. PhD thesis, Universite P. et M. Curie, Paris, France, 2000. under review.
STRUCTURAL CONTROL OF BRIDGES: RECENT EXPERIENCES IN CABLE-STAYED BRIDGES JOAN R. CASAS Escuela Tecnica Superior de Ingenieros de Caminos, Canales y Puertos (UPC) Gran Capitan s/n. Modulo CI. 08034 Barcelona, Spain E-mail:
[email protected] The paper presents different case studies on dynamic testing of cable-stayed bridges as the first necessary step to develop afterwards an efficient strategy for the structural control of such important structures. The model identification process and the calibration of the most accurate dynamic model to predict the actual real behavior of the bridges are presented. Based on the feasibility of obtaining an up-dated calibrated model of the performance of such complex structures, as shown in the paper through 3 case studies, the last part of the presentation is devoted to the most appropriate techniques of active structural control to be applied in those structures taking into account the presence of cables in the structure itself. Finally, a theoretical example of the application of a decentralized strategy of structural control is applied to an existing bridge subjected to earthquake excitation. Prospective applications of long-term monitoring to develop in the near future smart cable-stayed bridges are also presented.
1
Introduction
A correct and efficient strategy of active control in a bridge is feasible only if an updated calibrated dynamic model of the real bridge is available. This may be difficult to obtain using only the theoretical data and information from the existing drawings. Although these general considerations are valid for most structures, they are especially relevant in the case of bridges, partly due to the usual lack of nonstructural elements that may affect the actual (measured) stiffness and damping, as compared with the same properties computed with numerical or analytical models. In the case of complex structures as cable-stayedbridges this difficulty becomes still more evident. However, the use of dynamic tests has been recognised as an effective method for data collection of mass, stiffness and damping properties. The results from a dynamic tests may be very usefull to up-date and calibrate a dynamic model of the real constructed bridge. This will be shown in the following 3 case studies of dynamic tests performed in cable-stayed bridges. 2
Alamillo bridge
The Alamillo bridge is one of the seven bridges built in Sevilla (Spain) because of the Universal Exhibition EXPO-92 (figure 1). The deck of the bridge (200 m span) is a hexagonal steel box girder with 4.40 m depth. Every 4 m two lateral cantilevers 13.20 m in width formed of steel ribs support a reinforced concrete slab of 23 cm thickness forming the traffic carriageway. The pedestrians pass over the upper flange of the steel box. The deck is connected to the pylon by 13 pairs of parallel stays. The pylon is a composite (steel-concrete) structure. The height of the pylon is 85
86 134.25 m. The pylon has an inclination- of 32° to the vertical, which makes it possible to balance the forces in the cable stays without the use of back-stays. A description of the bridge and the construction process can be found in [1]. A dynamic test was performed in the bridge to determine the modal properties to validate the mathematical and full-aerolastic models used in the design [2,3]. The main results concerning the dynamic parameters of the bridge and their comparison with theoretical ones are summarized in table 1. As deduced from the table, the agreement between dynamic parameters of the real bridge and theoretical and scaled models (tested in wind-tunnel) was completely satisfactory.
Figure 1: Alamillo bridge (Sevilla, Spain) Table 1. Natural frequencies in the Alamillo bridge
Vibration mode
| |
3
Transverse pylon 1 Longitudinal (pylon + Longitudinal (pylon + Transverse deck 1 Longitudinal (pylon + Torsion deck 1 Transverse pylon 2 Longitudinal (pylon + Torsion deck 2 Longitudinal (pylon + Transverse deck 2
Theoretica f(Hz) 0.292 0.373 0.610 1.088 1.191 1.235 1.583 2.196 2.298 2.312 3.244
Aerolasti f(Hz) 0.30 0.39 0.65 1.20 1.19 1.11 1.67 1.97 2.19 — 3.4
Actual f(Hz) 0.30 0.40 0.66 — 1.205 1.155 1.537 2.155 2.295 2.78 —
Chaco-corrientes bridge over the Parana River (Argentina)
The Chaco-Corrientes Bridge consists of two independent half-bridges with prestressed concrete deck and two main longitudinal box girders made of precast segments. It has a central free span of 245 m between pylon axes and lateral spans
87
of 112.5 m. (Fig. 2). The bridge was completed in 1973; during the first 10 years of service, the bridge vertical profile suffered significant creep deformations that caused deterioration/loss of the elastomeric supports of the simple supported spans connecting the half bridges, and produced a systematic slowdown of traffic on the bridge due to the loss of vertical alignment of the road surface.
Figure 2: Chaco-Corrientes Bridge (Corrientes, Argentina)
In order to compensate for the accumulated deformations, the bridge Owner/Operator decided to replace all original stays before formulating a programme for cable re-stressing to compensate for the creep deformations. To assist in that process, a series of ambient vibration tests and impulsive tests were performed in order to define with sufficient accuracy the degree 'of geometrical corrections that could be introduced while complying with structural safety requirements. The dynamic tests were judged to be a reliable method to determine the current deck stiffness, since the actual static displacements measured during replacement of the cables are highly sensitive to temperature gradients, and therefore are not sufficiently accurate for this purpose. A series of ambient vibration measurements were performed on the bridge at the sections where the groups of stays are fixed to the deck [4]. Typical Normalized Spectral Density Functions (NSDF) of the vertical component of accelerations are shown in Fig. 3 from which the natural frequencies in Table 2 are obtained. QJ X3
1000
^ E
aoo
TO
600
J ^
400 200
0.587 ' 2-404
0 -K^--^ 6
8
10
12
14
16
18
Frequency [Hz]
Figure 3: Typical Normalized Spectral Density Functions (NSDF)
20
88 Table 2. Measured Natural Frequencies by NSDF and Phase Dispersion Procedures
Frequency of picks 1 2 3 4 5 6 7 8 9
ANPSD [Hz] 0.567 1.621 2.404 4.058 11.810 13.013 -
Phase Dispersion [Hz] 0.540 1.648 2.384 2.948 4.043 9.208 11.826 13.104 14.173
The cable forces due to permanent loads were also measured indirectly by means of records of transverse vibrations of the cables. The locked-type cables have low bending stiffness in relation to their mass and axial force, and consequently present little dispersion of the transverse waves. In this way an up-dated model of the cables was obtained too. 4
Zarate-Brazo Largo bridges over the Parana river (Argentina)
As depicted in Fig. 4 they are two almost identical cable-stayed bridges with four roadway lanes and one railway line on a steel deck (330 m central span and 110 m lateral spans) that conform together with the access viaducts, the so-called ZarateBrazo Largo Road/Rail Complex. One of the cables failed unexpectedly due to the combination of factors that had been ignored due to lack of proper maintenance, under permanent and normal traffic loads and normal meteorological conditions. The main reason behind the sudden rupture was the important degree of corrosion of the parallel wire cables mainly close to the bottom anchorages.
Figure 4: Cable-stayed part of the Zarate-Brazo Largo Bridge Complex (Argentina)
89
The objective of measurements of deck vibrations is to determine through tests the most significant natural frequencies of the cable stayed bridges of ZBL in order to calibrate the numerical models used in the evaluation of traffic and wind effects. Table 3 summarizes the main natural frequencies involving vertical components of displacement. In summary, the three lowest frequencies of the dominant modes of both bridges that involve vertical deck vibrations of the two ZBL bridges are: 0.44 Hz, 0.58 Hz and 0.63 Hz Table 3. Natural frequencies of ZBL Brigde [Hz]
Parana de las Parana Guazii Bridge Palmas Bridge 0.441 0.436 1 0.584 0.572 2 0.641 0.632 3 0.688 0.672 4 0.762 0.749 5 0.930 0.914 6 0.968 0.969 7 1.003 0.994 8 1.074 1.072 9 1.111 1.112 10 1.153 1.160 11 1.223 1.221 12 1.270 1.267 13 1.352 1.345 14 1.393 1.399 15 1.508 1.496 16 These measured frequencies provided useful information in two respects: i) The structural behaviour of the bridge. The new set of measurements confirm that the behaviour of the two bridges is in agreement with the numerical model and in close agreement with measurements performed at the end of construction. ii) Provided actual field data to calibrate the numerical model used to assess the structural safety of the bridges (deck, cables and pylons) after 22 years of service. Mode
A set of dynamic tests in the cables were also performed to provide a complete and updated set of natural periods and cable forces under permanent loads of all cables to serve as reference values and to provide the "initial" forces for cable replacement operations in the future, and also to help identify possible changes of cable stiffness with respect to the reference state at the end of construction (1977). The anchor pendulums are steel trust that link the deck with the anchor pier of the bridge. To evaluate their actual condition and effectiveness in anchoring the deck at this point, a set of measurements of the vertical displacement at their location were
90 carried out. This is a very important point because the incorrect performance of the anchor pendulums derives in the modification of the boundary conditions at the end of the bridge, and this may affect in a very relevant way the dynamic behaviour of the bridge. It has important influence in updating the theoretical model too. The relative vertical displacements between the top of the pendulums and the top of the anchor piers was recorded under normal traffic conditions. These records show variations of vertical displacements caused by traffic loads with changing signs depending on the instantaneous position of the moving load. For the purpose of comparing the experimental results with predictions by the numerical model, the influence line of the reaction at the pendulum was computed with the model of the bridge, and superimposed in scaled form with typical displacement records generated by a single truck moving along the bridge. These results are shown in Fig. 5. The pendulum displacements differ significantly from the computed influence line. Therefore the assumption of fixed zero displacement is not valid.
0
50
100
150
200
250
300
350
400
Distance [m]
Figure 5: Theoretical and measured influence lines of the force on roadway pendulum
5
Application to active control of cable-stayed bridges
We have shown how the calibration of the theoretical dynamic model for a cablestayed bridge can be performed via dynamic testing. This makes possible to get a very accurate model to predict the real dynamic behavior of the bridge. Thus, the techniques of active control can be used in a more efficient way. In the following we will see an application of active control in a cable-stayed bridge. Although the active control is only theoretical and simulated, the bridge is real. The idea is to design an automatic control system for a cable-stayed bridge to actively counteract external environmental forces generated by earthquakes. The introduction of this intelligent feature in the structure turns an otherwise passive bridge into a smart bridge, capable of reducing the amplitude of the deflections of
91
its deck down to a desired level automatically. The control strategy proposed here uses a subset of the stay cables as active tendons to provide control forces through appropriate actuators. Each individual actuator is controlled by a decentralized controller that only uses local linear velocity and local linear relative displacement information. The effectiveness of the control algorithm is tested on a threedimensional model of the Quincy Bayview bridge [5,6]. The length of the main span is 297 m, and the length of each of the side spans is 145 m. The width of the deck is 13.2 m (from cable center to cable center). The bridge has a total of 56 stay cables. This paper uses the same bridge modeling philosophy outlined in [5,6]. The values of the modal frequencies found with the finite element model developed here differ from those found in [5,6] very slightly. The computed mode shapes, however, are almost the same for the first 20 modes. Therefore, we can conclude that we have a very accurate model of the bridge, as this has been compared with the results coming from a dynamic test in the real bridge. The finite element model uses 213 three-dimensional beam elements (six degrees of freedom per node) for the deck and towers, and 56 one-dimensional truss elements. The first 10 modal frequencies (in Hz) computed from the finite element dynamic model of the bridge are the following: 0.38, 0.45,0.52, 0.55, 0.60, 0.65, 0.70, 0.74, 0.80, 0.82. Before to proceed with the controller design, we need to assess the uncontrolled bridge performance in the presence of seismic excitation. Let such an excitation be provided by the corrected vertical acceleration record of the Taft Earthquake. Most of its energy is concentrated in the range of 0 to 13 Hz. This is a good test signal to assess uncontrolled performance, since it can provide a high enough level of excitation to practically every on of the lower vertical modes of the bridge. Figure 6 shows the bridge's deck vertical deflections at t = 2 and 10 seconds. The maximum deck vertical deflection is 0.18 meters. After observing the vertical deflections of the deck of the uncontrolled bridge caused by the earthquake, it was decided to place 6 couples of active controllers. The design of the decentralized active control system is fully explained in [7].
''
j
A^
i
i
&
100
200 300 400 posfem ikMgbndgc deck n m a t r s
OTffl ICO
299 300 400 postisHt itoog bridjp deck ift i w t m
Figure 6 : Vertical deflections along span for t= 2 s (left) and t = 10 s (right)
$00
92
Then the controlled response of the bridge in the presence of the same seismic input (Taft earthquake vertical acceleration component) is shown in Figure 7. In this case, the maximum deflection is only 0.016 m. Although only local information has been used to control the entire bridge, the deflections have been reduced substantially.
tinrc * i seconds
posiisa afeng bridsss l , 0 < m < l . With the assumption of zero-order hold, the solution of Eq. (1) in discrete-time format is x(fc + 1) = [A]x{k) + [£i]u(fc - I) + [B a ]u(* - / + !) + [E]w(k)
(2)
where [A] = e^T,[B1] = [AC}-\[A] - [A}™)[BC], [B2] = [AC]^([A]™ [I])[BC], and [E] = [Ac]'1 ([A] - [I\)[EC]. In direct output feedback, the control forces are calculated directly from the multiplication of delayed output measurements by constant feedback gains. Thus, u(fc — £) = [G][D]x(k — (.). In order to represent Eq. (2) as the traditional first-order difference state equation, we need to define a new augmented state vector x(fc) which includes the control force vectors from u(k — 1) to u(fc — t). Then the new first-order difference equation will be x(fc + 1) = [A]x(k) + [B]u(fc) + [E]w(k)
(3)
where
[A] =
[A] [Bi] [B2] 0 0 0 -[I] 0 0 0 0 m
• •
• o•
0
• •
0
• o
, [B] =
0 0
0 0
0 0
0 0
• • [T\ . . 0 .
\[E]
•
0 0
0 0 , [E] =
0
Lm-J
(4)
0 . 0
with the modified output matrix [D] = [ [D] 0 0 ... 0]. For a controllable and observable system, there exists an optimal output feedback gain matrix [G] such that the discrete quadratic performance index oo
J = £{x T (fc)[Q]x(fc) + uT(fc)[iqu(fc)}
(5)
fc=0
is minimized subject to the system constraint. In Eq. (5), [Q] is a positive semi-definite augmented response weighting matrix constructed from [Q] and [R] is a positive definite control force weighting matrix. If the control system
96
performs well under random excitation, it will work equally well for random initial conditions, and vice versa. Consider the system under random initial disturbance XQ = [ x(0) 0 0 ... 0 ] , the discrete-time optimal output feedback gain matrix [G] satisfies the following simultaneous linear algebraic equations 4 , 5 : ([A] + [B][G][D])T[H]{[A] + [B][G][D]) - [H] + {[Q] + [D]T[G]T[R][G\[D]) = 0 (6) {[A] + [B][G][D])[L]([A] + [B][G]{D])T - [L] + [Xo] = 0 (7) [5]T[S]([A] + [B][G][D])T[L][D]T + [R][G][D][L][Df = 0 (8) where [H] = £ ~ o { [ ( M + [B}[G][D})T([Q] + [D]TMT[K\[G\[D]){[A] -T [B][G][D])k}, [L] is the Lagrangian multiplier matrix, and [X0]
+
x0 • x 0 .
2.2
Numerical Model and Analytical
Simulation
For a structure with a hybrid mass damper under earthquake loading xg(t) and time delayed active control force u(t-£<j) as shown in Fig. (2), the equation of motion of the system can be represented as XBr(t) XDB(t)
+
2£BUJB
—2/J,£DU>D
{ XBr(t) 1
0
2/^DWD
1 XDB(t) j
/ XBr(t) \ _
^i
u(t - td) +
w% ftu2D -1 -A*
,(*)
(9)
where XBT and XDB are the relative displacement of the system and the relative displacement of the damper with respect to the system, respectively. The parameters of the system 7 are given in Table(l). If we choose UD — 0 and £D — 0, the HMD system becomes an AMD system and if you force u = 0 it becomes a PTMD system. The pattern of sensor types are symbolized by X and V. For example, XDVD indicates that the relative displacement (XDB) and velocity {XDB) of the damper with respect to the system are measured. A study of the theoretical development and numerical simulation results indicates that the XDVD feedback in both HMD and AMD systems is more efficient and acceptable. The active damping ratio £c of AMD system is more sensitive to time delay than the HMD system. Thus AMD XDVD control type will be used in this paper to illustrate the required implemental procedures. The influence of delay time to the active modal frequencies u_c and active damping ratios £c of the controlled system is shown in Fig. (3). From previous
97
experimental results 8 , the calculated delay time of an active-tendon system is close to 40 msec. In order to emulate this real configuration of experimental setup, we set the delay time t d = 40 msec. The corresponding continuous-time modal properties when ta= 0 and 40 msec ( A M D C - T ) are calculated by using the explicit stability formulas proposed by the authors 2 . The discrete-time control gains are calculated based on the same choice of weighting matrices [Q] and [R] as in the continuous-time. Since the real experimental requirement of sampling period is 22 msec, we choose the sampling period T= 20 msec in Eq. (2) to find the discrete-time output feedback optimal control gains. The control effects on modal properties denoted as A M D D _ T are shown in Table (2). In practical implementation, the control gains are saved inside the control program for finding the required control force. The modal damping ratios of A M D C - T ( O ) are reduced dramatically to those of AMDc-x(40) due to the 40 msec time delay. This effect can also be observed if we use the continuous-time gains in the discrete-time modal analysis with sampling period T=20 msec. The mismatch of modal properties between continuous-time and discrete-time is also due to the choice of sampling period. The continuous-time case with T -> 0 is impossible to achieve in real implementation. So the discrete-time modal properties also provide a tool to predict the effectiveness of control under certain sampling period and delay time. If we use the discrete-time optimal gains for the AMD system, A M D , D _ T ( 4 0 ) , the control effect is even better than the ideal continuous-time case without delay. As we check the difference between the control gains, the time delay and sampling period effect had been automatically included inside the optimization process. There is no phase-shift needed for the output feedback control algorithm, and actually it is also difficult to perform the phase-shift modification. There is another important issue that needs to be clarified before we implement the proposed algorithm in real-time application. According to the derivation of discrete-time control gains, the sampling period of applying control force, Tu, should be the same as the sampling period of response measurement, T. In order to predict the response of a continuous-time model correctly, however, we need to simulate the response of the corresponding discrete-time model in a smaller sampling period. From the verification testing, the sampling period of the Real-Time Structural Simulator is chosen as 1.82 msec. The corresponding response of uncontrolled and PTMD systems are verified to be the same as those in the continuous-time approach (Table(4)). The 1940 El Centro earthquake (N-S component) in Fig. (5) is used as the base excitation to verify the control effect and to find the required control forces. The discrete-time analytical solutions are simulated by using Tu— 22 msec and T= 1.82 msec.
98
3 3.1
Experimental Verification Experimental Setup
The integrated pseudo real-time verification system as shown in Fig. (4) is constructed to perform the pre-implementation testing that includes two parts as follows: D S P Controller System: The required control force is calculated by a dedicated controller PC which is equipped with a QPC/C40B TIM Carrier Board with one TMS320C40 processor, a 12-bit A/D Converter Board, and a 12-bit D/A Converter Board. The TMS320C40 is the first of Texas Instruments 'TMS320C4X' generation of floating-point processors that is capable of performing the Programmable Erasable Read-Only Memory (PEROM) function. The Read-Only Memory (ROM) is used to store the programs that are executed repeatedly. This set of data acquisition system is also equipped with a DSPLINK digital system expansion interface, which is a high-speed, bidirectional bus that allows input/output directly to/from the DSP between A/D or D/A converters, without using the I/O bus on the host machine. The proposed control strategy is coded in C / C + + language and is downloaded to the PEROM of DSP to perform independent data acquisition and control force calculation. Real-Time Structural Simulator: The digital structural simulator employed in this experiment utilizes a DT-2801 data acquisition board made by Data Translation Inc. It is configured on a board that plugs into an ISA slut in a PC's expansion bus and features an on-board microprocessor and 8 differential 12-bit A/D input channels and two 12-bit D/A output channels. The actual real-time response is performed by a customer-written program downloaded to the microprocessor on the DT-2801 board. This simulator software is written in C language in order to access the hardware and communicate with A/D and D/A converters. Although the discrete-time model will converge to continuous-time model if the sampling period is close to zero, the time needed to calculate the analytical response of the discrete-time model inside the microprocessor should also take into consideration. The sampling period of the Real-Time Structural Simulator can't be too small in order to provide enough time to perform the calculation. The real sampling period will depend on the algorithm of the program and the numbers of data needed to be recorded. It is thus very important to identify the sampling period with respect to different codes.
99
Real-Time Control Process: To perform the real-time control, the output of the system at each time step is measured by DSP Controller and the required control force is then calculated inside the DSP and send out the analog voltage signal through ADC to actuator to apply the appropriate control force to the system. In this experiment, this control force analog voltage is routed directly into the Real-Time Structural Simulator to perform the pseudo realtime testing. In order to emulate the 40 msec delay time, we set the internal clock inside DSP program as 0.4 kHz (2.5 msec) and the resulting sampling period for response measuring T and control force application Tu is 22 msec. The required time to calculate the control force is also identified as 44 msec. Detail descriptions can be found in Table (3). The internal clock inside RealTime Structural Simulator is set as 1 kHz (1 msec) and the real sampling period to perform the response of the HMD benchmark model is identified as 1.82 msec. This choice of sampling period is verified by experiment to be small enough to emulate the continuous-time response of the HMD benchmark system.
3.2
Experimental
Verification
According to the estimated maximum response, we then setup the scaling factors to conduct the experimental verification as shown in Table (4). If any of the signals either inside the DSP Controller or the Real-Time Structural Simulator exceeds the preset limitation, the program will be terminated automatically to protect the system from saturation which may cause damage in real application. The control effects of the ideal continuous-time control without time delay are compared with the discrete-time analytical solutions and the experimental measurements from the Real-Time Structural Simulator as illustrated in Table(5). Without the action of control force, the experimental results show that the Real-Time Structural Simulator can emulate the response of uncontrolled and PTMD systems very well. And the results of discrete-time analytical analysis are very close to those from real measurements of the experiments. The only exception is the A M D C _ T ( 4 0 ) case, which uses the continuous-time gains inside the implementation of the discrete-time controller. Since the response of this system has exceeded the hardware limitation of the Real-Time Structural Simulator, no further experiments can be conducted unless the preset scaling factors are increased. With the XDVD output feedback type, the required stroke measurements of system with 40 msec delay time (denoted as AMD (40)) are compared with PTMD in Fig. (6). The measurements of absolute acceleration and relative displacement of the
100
AMD system with 40 msec delay time are depicted in Fig. (7). The corresponding control forces calculated by analytical simulation and experimental measurement are compared in Fig. (8). 4
Conclusions
The issues along with the implementional procedures include discrete-time control, time delay, model construction and experimental verification are demonstrated in this paper. The control effects on the HMD benchmark system with time delayed control forces have been investigated thoroughly by both analytical and experimental verification. The proposed discrete-time output feedback optimal control algorithm is proved to provide the desired control effect and stability of the controlled system is also guaranteed. Moreover, the integrated real-time simulation system provides an economic and efficient testing platform. It can also be used to conduct reliability analysis of the integrated system with parameter uncertainty or measurement noise. Acknowledgments This research is supported in part by the National Science Foundation under grant No.CMS9402196 and by the Multidisciplinary Center for Earthquake Engineering Research under grant No.MCEER-992401. References 1. K.J. Astrom and B. Wittenmark, Computer-Controlled Systems: Theory and Design, Prentice Hall, Upper Saddle River, New Jersey, 1997. 2. S.Y. Chu and T. T. Soong. "Time Delay Effect on Direct Output Feedback Controlled Mass Damper Systems", to appear in Proceedings of the 2000 American Control Conference, 2000. 3. S.Y. Chu, T. T. Soong and A.M. Reinhorn. "Integration Issues in Implementation of Active Control Systems", to appear in Proceedings of the Second European Conference on Structural Control, 2000. 4. L.L. Chung, C.C. Lin and S.Y. Chu, "Optimal Direct Output Feedback of Structural Control", J. Engrg. Mech., ASCE, 119, 2157-2173, 1993. 5. L.L. Chung, C.C. Lin and K.H. Liu, "Time-Delay Control of Structure", Earthquake Engrg. and Struct. Dynamics, 24(5), 687-701, 1995. 6. Gene F. Franklin, J. David Powell, and Michael L. Workman, Digital Control of Dynamic Systems, Addison-Wesley Publishing Company, Inc., 1990.
101
7. C.C. Lin, J.F. Wang and Y.C. Gau, "System Identification and Vibration Control of Structures with Tuned Mass Dampers", Proc, 3rd R. 0. C. and Japan Joint Seminar on National Hazards Mitigation, 436-450, 1993. 8. S. McGreevy, T.T. Soong and A.M. Reinhorn, "An Experimental Study of Time Delay Compensation in Active Structural Control", Proc. SEM 6th Modal Analy. Conf., II, 733-739,1987. 9. T.T. Soong, Active Structural Control: Theory and Practice, John Wiley k Sons, Inc., New York, 1990. Table 1. Parameters of the System
System parameters Mass , TUB (kg) N a t u r a l frequency , UJB ( H Z . ) Damping ratio , £g (%) D a m p e r mass ratio , /x = mo/ms P T M D , D a m p e r designed frequency , WB ( H Z . ) P T M D , D a m p e r designed damping ratio , £ D (%)
Parameter values 6897.5 1.00 2.00 0.05 0.93 11.00
Table 2. Control results on modal properties
Cases
td (msec)
[-
N/cm
[G\ - N • sec/cm]
Uncontrolled PTMD
(Hz) 1.00 0.87 1.07
(%) 2.00 6.90 6.31
0.87 1.03 '0.86' 1.08
13.15 7.87 '4.77' 4.49
0.86 1.08 '0.87' 1.03
3.02 3.09 15.87 8.74
Continuous-Time AMDC-T
0
[0
-109.70
0
-7.25]
AMDC-T
40
[0
-109.70
0
-7.25]
0
-7.25]
Discrete-Time with T = 20 msec AMDC-T
40
[0
-109.70
AMDD_T
40
[0
-92.09
0
-12.30]
102 Table 3. Distribution of microcode execution time
Time Description (msec) 40 Target emulated delay time 2.5 Clock-Setup referred to 4 A/D's conversions in parallel 20 Conversion time for all channels (32 ch) on the board / 1 cycle 2 Program execution time / 1 cycle 22 System total sampling time / 1 cycle 44** Active force application delay time ** Achieved delay time needs 2 cycles of program computation
Table 4. Setup of Experimental Scaling Factors
Software/Hardware HMD Model Structural Simulator DSP Controller
yU'lmax
max
max
±30.48 cm (±12 in) ±10 Vdc ±2.5 Vdc
±203.2 cm/sec (±80 in/sec) ±10 Vdc ±2.5 Vdc
±4450 N (±1000 lb) ±10 Vdc ±10 Vdc
Table 5. Maximum response under El Centro earthquake
Cases
(XBr) (cm)max Continuous-Time Analytical Solutici n 15.16 Uncontrolled 0.61 PTMD 0.36 8.76 0.40 9.96 AMDC-T(O) Discrete-Time Analytical Solution 15.24 0.61 Uncontrolled PTMD 0.36 8.76 11.21 0.47 AMDC-T(40) 9.72 0.40 AMD£,_ T (40) Experimental Measurements Uncontrolled 0.61 15.23 PTMD 8.76 0.36 N/A N/A AMDC-T(40) AMD£,_T(40) 0.41 10.15 max
isfs)
max
y^Jmax
(cm)
(N)
29.67 22.58
2687.8
29.76 43.44 23.21
5164.41 2781.4
29.61 N/A 22.72
N/A 2707.37
103
Digital Control System (DCS) Control Command Calculator (CCC) System Information Control Command (Digital) (Digital) j Data Acquisition/Conversion System (DACS) ADC DAC
Control Command (Analog),
Analog Control System
System Information (Analog)
Custom-Designed Signal Interface System (CDSIS) , pontrol Command (Analog)
Status Check
Active Control Force Generation System (ACFGS)
Measuring Equipment (Sensors)
Displacement, Velocity, Acceleration, Force
Active Control Force
Excitation
Signals (Analog)
Response
Structure
Real-Time Structure or Structural Simulator Figure 1. The General Hardware Function of Active Control Systems
.
XDB
—•
, ITlD
u —
KD^W
^ J
•
CD XBr
ITIB
K
•
^3c» ^»
Xg
Figure 2. Schematic diagram of hybrid mass damper system
104 1.1
: • .
: —
-
"
"
^
14
:
12
'
1.05
l
I
-
*«*
10
~ ' ~" - -: ^Sv.
u_p 6
3 0.95
•OC ;
4 0.9
2
7s^
0
0.85 0
0.02 0.04 0.06 Delay Time t (sec)
0.08
0
0.02 0.04 0.06 Delay Time t (sec)
0.C
Figure 3. Time delay effect on active modal frequencies and active damping ratios of AMD XDVD system 2
Real-Time Structural Simulator
DSP Controller System
E M L-J
l
Keyboard
Ti
Monitor
DT2801 Data Acquisition Board
L.
DSP LINK QI' c ' c * )B ' rIMCi " ,ie, ' B, > ard
Vjgg_nimn fl • DDDOD I-!
TooO o o
DDDOD
B
Kevboanl
16 Caunel D/A Converter
Custom-Design Signal Interface System (CDSIS)
Figure 4. The Experimental Setup of the Pseudo Eeal-Time Verification Test System
Figure 5. 1940 El Centro earthquake
105 30
'if
20
^
t:
4 11 }
10
, F 11111IfLK »V\;iK.iA: A A
0
vi i
^.-10 -20 -30
ZTMD
-
,•/ \•,A• i X A
" y i/ v tj v v \] \f
j
A A.
» j.
;
. ; / •
v
i * y '1 K H
1
i
10
15 Time (sec)
i
20
25
30
Figure 6. Stroke of damper with P T M D and AMD(40) (Experiment)
15 Time (sec) Figure 7. Absolute acceleration and relative displacement with AMD(40) (Experiment)
Figure 8. Control force with AMD(40) (Analytical vs. Experimental)
MONITORING OF INFRASTRUCTURES IN THE MARINE ENVIRONMENT
A.
DEL GROSSO
Department of Structural and Geotechnical Engineering, University of Genoa, Via Montallegro, 1 -116145 Genoa, ITALY e-mail: delgrosso@diseg. unige. it Maintenance, implemented through inspection and rehabilitation cycles, of infrastructures has taken a paramount role in modern civil engineering research because of the very large amount of aged infrastructures existing in the world. Among these infrastructures, those interacting with the marine environment, i.e. port and coastal structures, structures for oil and gas production and transportation, are posing to the engineering community very interesting and challenging problems. After a review of the maintenance strategies more useful for engineered structures in the marine environment, the paper addresses some topics of the longterm monitoring of breakwaters, internal port structures, offshore platforms and submarine pipelines, showing the interdisciplinary aspects arising from these new fields of the civil engineering activity.
1
Introduction
Looking at the scientific and technical journals belonging to the field of civil engineering over the last 20 years, we observe that many new and unconventional problems have been brought to the attention of the engineering community. Moreover, we will also discover that many prestigious journals, well known in other disciplines, are publishing papers somehow related to civil engineering problems. The rationale staying behind this observation, lies in the fact that managing constructed facilities is becoming more critical than building new ones. Economical and social reasons, sometimes more stringent than purely technical reasons, stay underneath this fact, but by no way we could avoid to face the challenges of this highly interdisciplinary context. Especially infrastructure systems form the object of these new developments, because on the one hand they are more subjected to obsolescence and, one the other hand, they involve economical and financial aspects more important than other engineered structures. In addition to aging and physical degradation, infrastructure obsolescence may have different sources. According to Lemer [1], the factors that can cause obsolescence can be grouped into four categories: 1. technological changes, that influence the scope or level of services the infrastructure is to provide; 2. regulatory changes, that impose new requirements on infrastructure; 3. economic or social changes in the market within a region, that alter the demand placed on infrastructure; 107
108
4.
changes in value or behavior of the people who use and own the infrastructure, that similarly alter demands. System management is therefore becoming at least as important as system design in the continuous balancing of safety, usability and profitability of an infrastructure. Such balance introduces the concept of fitness for use. A dynamic measure of the fitness for use can form the basis for decision making by expert panels in establishing if an infrastructure needs intervention. Relative measures of fitness for use can also be very useful in establishing ranking systems, able to help decision makers, responsible for the management of large infrastructure systems, in optimizing investments for repair or replacement. In the simplest form, fitness for use can be expressed by a weighted summation of scores, attributed to the infrastructure according to different criteria: Fu = I ; Wi ^ ,
(1)
where Sj is the score and w( is the corresponding normalized weighting function, such that Zi Wj = 1. However, multiplicative models and different multi-criteria as well as single-criteria measures of the fitness for use can also be introduced, and qualitative methods as well as quantitative, based on complex mathematical models, can be used to perform scoring. Monitoring of an infrastructure involves periodical reassessment of the fitness for use, in order to determine its decay in time and, whenever the measure indicates that a minimum acceptable limit is being reached, refurbishment activities shall be planned. Infrastructures interacting with the marine environment, like ports and facilities for oil and gas exploitation in the continental shelf, represent a very significant casestudy for the above principles, because of their extreme economical importance and because of the concern about the effects that these facilities may produce on the natural environment. Fixed marine structures are primarily designed to withstand environmental forces. Knowledge about phenomena related to environmental dynamics has greatly improved in the recent years, and today's standard design loads are substantially increased with respect to past standards (e.g.: the return period of the design wave has been increased in API standards from 25 to 100 years after the hurricanes in the Gulf of Mexico of 1964 and 1965 [2]). Fatigue, corrosion and material degradation due to chemical and biological phenomena are more severe in the marine than in any other environment. Many coastal structures are inherently subjected to damaging by the sea (rubble-mound breakwaters) or are definitely conceived for being modeled by the waves (berm breakwaters). In addition, interaction with the environment may cause effects that in turn can modify actions on structures ( e.g.: reflected and refracted waves, drag forces, scouring, interaction between sea floor and pipelines, etc.). All the above considerations indicate that life-cycle management of infrastructures in the marine environment cannot be conceived without putting in
109 place fitness for use reassessment and requalification programs based on condition monitoring of structural components. This paper is intended to examine some of the aspects related to structural monitoring in this context, with the aim of calling to evidence the needs for developing and testing innovative technologies. Since monitoring is strictly related to maintenance strategies, a brief summary of the maintenance approaches suitable for application to infrastructures in the marine environment will be presented first. 2
Maintenance strategies
The definition of a maintenance policy shall be a part of the design process, in order to keep costs and benefits into a constant equilibrium during the entire lifetime cycle of an infrastructure. It is noted that, sometimes, infrastructures have very long design lifetime (50 to 70 years or more). The concept linking design, maintenance and risk can be expressed through the minimization of the lifetime cost [3,4]: Qifetae = I + PV(M) + PV(R) + PV(pf Cf)
(2)
where : I is the investment (construction) cost, M is the cost of monitoring, R is the cost of repair, p f is the annual probability of failure, Cf is the cost for the consequences of a failure, and PV is the present value operator. The last term of equation (2) is clearly a measure of the risk or a cost associated to failures. Monitoring techniques are defined in terms of the objectives that monitoring activity should obtain in function of the maintenance strategy adopted. Usually, a distinction is made between corrective and preventive maintenance. A corrective maintenance strategy is adopted when repair or rehabilitation is undertaken after a failure has occurred. This strategy may be suitable only for cases in which the global cost associated to the failure is low. In this case, the role of monitoring is just to detect if, when and where the failure has occurred. A preventive maintenance strategy, instead, is aimed at preventing the occurrence of failure states. Preventive maintenance is suitable when the cost associated with failure is high. Preventive maintenance strategies can further be classified into: 1. state (or condition) based maintenance, 2. time based maintenance, and 3. load based maintenance. State based maintenance is the most general approach. According to this approach, the role of monitoring is to allow detection and evaluation of the state of the system in the real-time, through the definition of an appropriate condition index or, if mathematical reliability models are available, of a reliability index. Several Repair, Evaluation, Maintenance, & Rehabilitation (REMR) programs based on
110
such concepts have been established worldwide by large-scale organizations managing marine infrastructure systems [ 5, 6, 7]. Time based maintenance is applicable when statistical knowledge allows the reliable definition of age-related hazard functions of some system component. In this way, detailed inspection or substitution of the components can be prescheduled. Monitoring is eventually required to improve knowledge on component performances. This method is normally applicable to widely used mechanical components. A load based maintenance strategy can be adopted when probability models are available, allowing to correlate the occurrence of a given intensity of load to the state of the system. Such strategy involves running of detailed inspections and rehabilitation works, only after a given phenomenon has occurred. In this case, the role of monitoring concentrates more on environmental phenomena than on the response of the system. Of course mixed strategies may be adopted for complex systems, especially when fitness for use derives from the superposition of the state of independent subsystems, with different characterization in terms of mechanical complexity and availability of probability models. In the selection of monitoring techniques for the different maintenance strategies, the following aspects should be considered: 1. cost of acquiring data; 2. data processing techniques, 3. information value of the data. The above discussion will be useful in the following paragraphs, where some of the problem related to monitoring of port structures, offshore platforms and pipelines will be addressed. It is anticipated that port structures will receive major consideration. 3
Monitoring of port structures
In port and harbor infrastructure it is necessary to make a distinction among external breakwater structures, internal structures and equipment. 3.1
Breakwater structures
In-service monitoring of breakwater structures may involve different aspects, depending on the type of breakwater and on what the monitoring is aimed at. The specification of in-service monitoring is due to the fact that in many cases monitoring during construction is aimed at optimizing the design/construction process. Construction of large breakwaters usually takes several years to be completed and may involve settlement of the sea bed, erosion by the waves and
111
streams, tilting of caissons etc., and requires a large amount of topographic and bathymetric surveys. In-service monitoring can be aimed at: a) improving knowledge on the site or monitor the effects of the structure on the environment; b) validating design hypotheses or applicable stardards; c) serving as a basis for maintenance operations. In the last case, the design of the monitoring system should be addressed to acquire information able to characterize the state of the entire structure and should fully comply with the criteria discussed in the previous paragraph. A further distinction can be made between vertical wall and rubble mound breakwaters. In the first case, the safety concept is similar to that of any other structure and damaging is not permitted, because it will lead to a major failure state. Monitoring usually involve determination of the water pressure on the vertical walls (to verify design hypotheses), dynamic response of the caissons to impacting waves, structural integrity, permanent settlement or rotations of the caissons. In the second case, damages can be permitted because they are repairable through maintenance and they do not necessarily mean failure. Monitoring shall therefore address the recognition of the damage state suffered by the jetty, in terms of fracture or displacement of their elements both underwater and above water (evolution of the profile), permanent displacements of the crest and characterization of the flow through the body of the breakwater. A common aspect is the monitoring of environmental conditions (wind, waves, streams, dynamics of nearby coastlines) and the evolution of the sea floor close to the breakwater foundation, because it may disclose the tendency to erosion, soil liquefaction or soil failures, causing collapse of the breakwater. It should be noted that the presence of a breakwater always alters the coastal equilibrium state. In both cases, a clear understanding of the potential failure mechanisms is crucial in designing a monitoring program. Important monitoring programs are currently under way in several countries. In the U.S., it is important to underline the activity of the Army Corps of Engineers, that is publishing very useful manuals and guidelines to design and run breakwater monitoring programs [8,9,10]. In addition, USACE is responsible for the conductance of large-scale monitoring of breakwaters and complete coastal engineering projects. A review of the most relevant experiences performed in the US is contained in the works edited by Magoon and Davidson [11]. In Europe, several projects have also been started. The largest and the most complete of them are the monitoring of the rubble mound breakwaters in Zeebrugge [ 12] and in the port of Sines [13 ]. A fairly small but still interesting program is being started in the Port of Genoa, where an old existing breakwater (Duca di Galliera) is being refurbished in order to recover the present damage state. According to the design, the jetty will be transformed into a berm breakwater [14]. It is planned to record the displacements
112
of the crest both during refurbishment and in the long-term, and to monitor the evolution of the critical profiles of the jetty. Monitoring for maintenance purposes is typical of rubble mound breakwaters. Among the different problems, two of them will be addressed in detail: a) detection of global displacements of the crest, and b) monitoring the evolution of the profiles of the jetty. 3.1.1
Displacement monitoring
The spatial distribution of settlement and rotations is of interest for the monitoring of breakwater behavior. Traditionally, displacement monitoring is performed by means of precision topographical surveys. However, modern technologies are available to perform this task and obtain data ready for computer processing. A readily available technology is represented by fixed GPS networks. A review of recent applications of this technique has been presented by Duff [15]. The method will be experimented on the Duca di Galliera Breakwater in the Port of Genoa. A future technology can be represented by satellite interferometric measurements, using fixed networks of corner reflectors [16]. 3.1.2
Profile evolution monitoring
This is the crucial aspect in the monitoring of breakwaters. An evaluation of the effectiveness of the different methods that can be used to this purpose is presented in [3]. Topographical, photogrammetric, and direct surveys using specially designed cranes have been used to monitor evolution of the emerged part of the jetty. Also, interesting applications of image recognition techniques have been used in South Africa by Hough and Phelp [17] to detect the displacement of the elements of a jetty from aerial images. For inspection and survey of the submerged part of the jetty, several traditional methods can be used, from diver's inspection, to side scan sonar bathymetry, etc. However, the interest is increasing towards emerging technologies such the use of ROV, multibeam Sonar and airborne Lidar systems. Especially, vessel operated multibeam Sonars are able to automatically produce accurate digital terrain models (DTM) of the submerged part [18]. Airborne Lidar Systems, such as SHOALS, are the most interesting technology because they can produce with a very reasonable accuracy a DTM of both the emerged and of the submerged parts, up to water depths close to 40 meters in optimal conditions. A review of the recent applications of the SHOALS system can be found in [19]. 3.2
Internal Structures
Internal structures in a port environment that may be subjected to monitoring are constituted by piers, docks, berths, dolphins, transportation facilities and buildings.
113
Such structures are not usually designed to withstand wave actions, but distress and degradation producing loss of functionality end even collapse may be caused, for example, by the following actions: a) Corrosion of steel and degradation of concrete and other materials; b) Excessive earth pressures due to overload on the quays; c) Overloading of slabs; d) Collision of Ships; e) Interaction with the sea floor (instability, erosion or dredging activities) f) External Hazards (earthquake, hurricanes, fire, etc). Corrosion and concrete degradation is one of the major causes of repair, especially after the diffusion of precast prestressed concrete elements that has taken place in the recent years for the construction of pier slabs, and because of the harsh microclimatic environment that sometimes is established between the slabs and the water surface. In a large port, monitoring by inspection usually involves important resources. Great interest is therefore devoted to instrumentation monitoring and some experiment is already going on. In the Port of Genoa, a 400 meter long gravity quay wall has been instrumented with fiber optic sensors to monitor the state of displacement potentially induced by nearby dredging activity [20]. Another application will be started soon, involving monitoring of a prestressed concrete pier slab with fiber optic deformation sensors and conventional sensors for moisture content, inserted in the beams at critical locations. It should be pointed out that fiber optic sensors look to be very suitable for application in this environment. Indeed, in addition to deformation sensing, fiber optic sensors can be developed for chemical, temperature, and other processes, thus allowing a high degree of integration of the data acquisition system. Stability, durability, independency on electrical disturbances, capability of being used in largely distributed networks render this technique superior to other sensing devices. 3.3
Equipment
A very large number of equipment involving important structural problems is used in ports. The degree of fitness for use of the entire infrastructure is largely dependent on the reliability and availability of such structures. Movable bridges, dock gates and locks, cranes and other loading/unloading equipment are among these structures. Structures for these types of application are usually steel structures, with the tendency in time to increase the slenderness (in order to reduce weight) and substitute welded box girders to traditional truss girders. As a consequence, instability to wind and lateral forces, insurgence of non-linear dynamic effects, excess of vibrations, fatigue cracking and steel corrosion are among the most frequently encountered problems for such structures, sometimes leading to
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unexpected collapses. In addition, many port facilities still make use of existing equipment well beyond their operating life. All such structures are usually subjected to detailed periodical inspection programs but, clearly, great benefits can be expected from instrumental health monitoring experiences. However, despite of the fact that modern equipment is normally operated through computerized systems, monitoring of the structural response is still far from practice. 4
Monitoring of structures for oil and gas exploitation
Structures for oil and gas exploitation are among the most technologically developed structures in the marine environment and, due the their economic and environmental impact, they have also been the first objective for the development of instrumental monitoring techniques. 4.1
Fixed offshore platforms and raisers
Very extensive research and application efforts have been devoted to structural monitoring of fixed offshore platforms. The number of major operating platforms is now about 6,000 worldwide, with approximately one third of them called upon for extended service or reuse [21]. Some oil companies are owning or operating hundreds of such structures. Monitoring and inspection programs also form the subject of API and ISO recommendations. A very large number of papers on the subject can be found in the literature, but a survey of them is out of the scope of this paper. Some basic references can be found in the mentioned works by Bea [21] and by Banon, Bea et al. [ 2]. For the interest of the present discussion, it will be pointed out that technologies such as GPS and fiber optic sensing can also be very useful in upgrading the instrumentation available for platform monitoring. It should be mentioned that, especially in deep waters, oil exploitation is also performed by means of moored or position controlled operating barges or semisubmersible structures or buoys, connected to the well head by means of steel risers. Although risers are connected to the barge and to the well head by means of special joints, they undergo severe stress states because of the relative movements between the barge and the well head. Design and control of risers is one the most interesting structural problems today arising from offshore technology. A very interesting monitoring application to workover risers has been proposed by Osen et al [22]. They have developed a system, based on strain monitoring at the two ends of the riser, able to advise the operating barge about the state of stress in the riser, and drive the repositioning of the barge in order to keep the stress state below allowable limits. In this application, LVDT sensor types are used to pick up
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strain data. However, Martinelli et al. [23] have shown that, in a similar application, arrays of fiber optic sensors have been able to work satisfactorily. 4.2
Pipelines
Development of marine pipelines for oil and, especially, for gas transportation has been very intensive in the recent years. Very large-scale projects have been realized and even larger projects are under study. It has been recognized [24] that hazards to underwater pipelines mainly (80 percent of reported failures) come from corrosion and third party external sources (navigation, fishing activity). Actually, inspection of existing pipelines is performed by the following tools: a) Intelligent pigs (internal inspection); b) Visual inspection by divers; c) ROV. Interest toward the use of continuous instrumental monitoring and new sensing technologies has been raised up by several parties, but no real applications are known at present. 4.3
Terminals
A very large number of open sea piers are existing worldwide to serve as terminals for oil export and import. Most of them are aged of more than 30 years, and quite a few built in the first decades of the 20th century are still in operation. This situation is posing a management problem to owners and public administrations because of the decision to close, substitute or refurbish the structures of these terminals that has to be taken in several cases. Usually, detailed periodical inspections are required to keep the terminals into operation. Interesting experiences, gathered by the California State Lands Commission have been presented in [25]. Besides these old or conventional types of structures, several projects are under way, aimed at realizing LNG terminals equipped with underwater storage facilities. Although even in this case practical applications are lacking, it is easily understood that instrumental monitoring and use of smart materials and sensors could greatly improve safety of such critical facilities. 5
Conclusions
Monitoring of infrastructures in the marine environment is a very broad subject and requires facing many different problems. From the cases discussed, it can be concluded that, in establishing monitoring programs for marine facilities, consideration shall be given, but shall not be limited to the following aspects:
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a) lifecycle cost optimization; b) engineering support to decision making, involving structural safety and reliability models, risk analysis and fitness for use models; c) interaction with the marine environment; d) sensing technology; e) data analysis and interpretation. The latter aspect is crucial when continuous automated instrumental monitoring is applied, because of the extremely great amount of information that is gathered from the structural response. This observation lead to individuate data analysis and interpretation as one of the key problems in structural health monitoring for complex infrastructure systems. References 1.
Lemer A. C , Infrastructure Obsolescence and Design Service Life. ASCE Journal of Infrastructure Systems 2 4 (1996) pp. 153-161 2. Banon, H., R. G. Bea, F. J. Bruen, C. A. Cornell, W. F. Krieger, D. A. Stewart, Assessing Fitness for Purpose of Offshore Platforms. I: Analytical Methods and Inspection. ASCE Journal of Structural Engineering 120 12 (1994) pp. 3595-3612 3. De Rouck J., K. De Winne, Full Scale Dynamic Load Monitoring of Rubble Mound Breakwaters. Report MAST2 Project MAS20023, University of Ghent (1994) 4. Moubray J., Reliability Centered Maintenance. Butterworth-Heinemann, Oxford (1991) 5. Chouinard L.E., G. R. Andersen, V. H. Torrey, Ranking Models Used for Condition Assessment of Civil Infrastructure Systems. ASCE Journal of Infrastructure Systems 2 1 (1996) pp. 23-29 6. Staneff S. T., C. W. Ibbs, R. G. Bea, Risk-Management System for Infrastructure-Condition Assessment. ASCE Journal of Infrastructure Systems 1 4 (1995) pp. 221-229 7. De Franco S., P. O'Connor, A. Tallin, F. Puskar, Development of a Risk Based Underwater Inspection Process for Prioritizing Inspections of Large Numbers of Platforms, paper OTC 10846, Offshore Technology Conference (1999) 8. USACE, Design of Breakwaters and Jetties. Publication No. EM 1110-22904, Chapter 12 Performance Monitoring Plan (1986) 9. USACE, Surveys of Coastal Structures. Publication No. CETN-III-41 (1991) 10. USACE, REMR Management Systems for Civil Works. REMR Technical notes OM-MS-1.1 11. Magoon O. T. , D. D. Davidson, Case Histories of the Design, Construction and Maintenance of Rubble Mound Breakwaters. ASCE (1995) 12. Van Damme L., J. De Rouck, Monitoring of Zeebrugge Breakwater. Proceedings Coastal Engineering (1998) pp. 1944-1956
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13. Pita C , F. Abecasis, A. Femandes, Sines Breakwaters Monitoring Program. In Reconstruction of the West Breakwater at Port Sines, Portugal, ASCE (1994) pp. 408-427 14. Fedolino M., P. Grimaldi, S. Stura, G.R. Tomasicchio, The Duca di Galliera Breakwater of the Port of Genoa, in Case Histories of the Design, Construction and Maintenance of Rubble Mound Breakwaters. ASCE (1995) pp. 16-22 15. Duff. K, GPS Monitoring and Control in the Marine Environment, Proc. Int. Conf. Monitoring and Control of Marine and Harbor Structures, Genoa (1999) 16. Dellepiane S., G. Bo, R. De Laurentiis, A. Del Grosso, Remote sensing and Coastal Monitoring: methods and applications. Proc. Int. Conf. Monitoring and Control of Marine and Harbor Structures, Genoa (1999) 17. Hough G., D. Phelp, Digital Image Processing Techniques for the Aerial field Monitoring of Harbor Breakwaters. Proceedings Coastal Engineering (1998) pp. 1787-1799 18. Prickett T., Coastal Structure Underwater Inspection Technologies. USACE Publ. CETN-III-62 (1996) 19. Irish J. L., J.K. McClung, W. J. Lillycrop, Airborn Lidar Bathymetry: the SHOALS System. International Navigation Association Bulletin 103 (2000) pp. 43-53 20. Del Grosso A., D. Inaudi, G. Brunetti, M. Fedolino, Monitoring of the San Giorgio Pier in the Port of Genoa with Fibre Optic Displacement Sensors. Proc. Int. Conf. Monitoring and Control of Marine and Harbor Structures, Genoa (1999) 21. Bea R. G., Reassessment and Requalification of Infrastructure: Application to Offshore Structures. ASCE Journal of Infrastructure Systems 2 2 (1996) pp. 45-53 22. Osen P., B. Johannessen, K. Stromsen, T.G. Werno, Instrumented Monitoring of Workover Risers, paper OTC 8797, Offshore Technology Conference (1998) 23. Martinelli M., A. Melloni, A. Gusmeroli, A. Tonini, P. Guaita, F. Trave, C. Barilla, C. Mariottini, G. Vegetti, C. Preti, G. Pagnoni e M. Pizzorno, Deployment of 32 FOIS Array for off-shore Structure Monitoring in the Adriatic sea, Proc. Int. Conf. Monitoring and Control of Marine and Harbor Structures, Genoa (1999) 24. Carpaneto R., E. Valente, Application of Innovative Monitoring and Sensing Technologies to Improve Safety of Marine Pipelines Against External Hazards. Proc. Int. Conf. Monitoring and Control of Marine and Harbor Structures, Genoa (1999) 25. Eskijian M.L., Structural Monitoring and Control of Marine Oil Terminals in California. Proc. Int. Conf. Monitoring and Control of Marine and Harbor Structures, Genoa (1999)
SEISMIC RETROFIT OF CHURCH BUILDINGS THROUGH BASE ISOLATION
A. DE LUCA, E. MELE AND C. GIAGNUOLO DAPS, Universita' di Napoli, Piazzale Tecchio n.80, Napoli ITALY E-mail:
[email protected] 80125,
Base Isolation has demonstrated to be a very efficient means for protecting buildings against destructive earthquakes. The application of base isolation to existing monumental buildings has been made quite extensively in the United States [1-9] to buildings charachterized by structural typologies which are not very close to the older typologies existing in Europe. In this paper the possibilities of application of Base Isolation to church buildings is demonstrated through dynamic numerical analyses which have modeled the entire building. The efficacy of Base Isolation is demonstrated through a comparison of performance with the existing monument and with the monument retrofitted via some techniques which are very common: insertion of rigid diaphragm. It is concluded that the insertion of rigid diaphragm does not always provide better performance and that base isolation could be adopted in most cases for the seismic retrofit of churches.
1
Introduction
Church buildings, which represent a large portion of the Italian monumental heritage, have frequently demonstrated to be very susceptible to damage and partial or total collapse in earthquakes. Masonry churches, in fact, have been designed and constructed for resisting with wide safety margin to vertical loads, but are not able to withstand horizontal actions, which give rise to tensile stress in the masonry elements. In this paper the seismic behaviour and a retrofit solution of a basilica type church (already studied in [10]) is investigated. The plan, facade and the main longitudinal and transversal sections of the church are provided in figure 1. In the schematic plan provided in figure 1 the main macro-elements, appointed respectively as L1-L3 for the longitudinal direction and T1-T5 for the transversal direction, can be derived. Such elements are very repetitive in the church building typology, thus the results of the analyses on this specific case study can be somehow extended to a number of structures having similar characteristics. 2
Structural analysis of the church
In this section the main results of the analyses performed on the structural complex of the church are provided. Dynamic analyses are carried out the by means of a FEM computer code, with reference to the EC8 elastic spectrum (PGA = 0.35g), with the
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120
aim of evaluating the principal modal shapes, the fundamental periods and the stress distribution in the single structural elements. The FEM model representing the structural system of the church is provided in figure 2. This model consists of 5941 joints 5572 shell elements and 56 frame elements used for timber trusses. PLAN SECTION A - A
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Figure 1. Plan, elevation and sections of the church studied
Figure 2. Finite Element Model of the church
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The analyses have evidenced that the fundamental periods of the church (figure 3 and table 1) in the transversal (0.449 sec) and longitudinal direction (0.337 sec) are sufficiently small to advise for a retrofit solution which makes use of base isolation.
\
Ti =0.449 sec
T 3 =0.337 s e c
T2=0.375 sec
Figure 3. Mode shapes and periods of the 3D FEM model. Fixed Base solution
LONGITUDINAL
TRANSVERSAL MODE 1 2 4 10 14 15 16 22 26 38
PERIOD (sec) M / M m (%) SM/M l o l (%) 0.46 0.377 0.293 0.273 0.208 0.189 0.183 0.164 0.153 0.084
42.852 6.3447 2.658 9.686 4.664 2.801 4.547 1.638 2.083 9.88
42.852 49.1967 51.8547 61.5407 66.2047 69.0057 73.5527 75.1907 77.2737 87.1537
MODE 3 9 13 25 27 29 32 39
PERIOD (sec)
M/M tot (%)
ZM/M l 0 , (%)
0.342 0.274 0.211 0.157 0.143 0.129 0.12 0.086
58.563 1.6353 6.6337 1.715 1.975 2.543 1.188 10.993
58.563 60.1983 66.832 68.547 70.522 73.065 74.253 85.246
Table 1. Modal participation factors: church without retrofitting
In figure 4 (a) and 4(b) are provided the schematic plans of the church at different levels. In particular in figure 4(a) the different zones in which rigid diaphragms can be inserted are appointed as A, B, C, D. In figure 4 (b) the location of BIS devices at foundation level are given.
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Figure 4. Plan of the church at different levels.
3
The BIS retrofit solution
The base isolation system has been designed to shift both the transversal and the longitudinal periods at 2.6 sec. The type of isolators to be inserted at the base of the church are HDRBs (high damping rubber bearings) having the characteristics given in figure 5.
J
t. r l
\ x Nssteel fe
plate
Rubber I
I
D
t r = rubber layer thickness H r =£t r = total rubber height D = diameter; A=(D/Hr)=3 +5 S = (D/4tr) > 30 av=3-10Mpa v = 10% Figure 5. Schematic representation and
main properties of BIS Devices
Given the position of the devices, which can be obtained from figure 4, and satisfying the requirements of figure 5, the dimensions for the isolators which are reported in table 2 have been obtained. Isolator D
K
n° mm mm
I I [
1 400 120
2 600 200
3 600 220
4 400 140
5 24 400 120
25 400 80
26-27-28 650 210
29-30 650 170
Table 2. Dimensions of BIS devices
In the FEM model of the B.I. structural complex a grid of very stiff beams has been introduced above the isolators. For the isolation system an equivalen linear viscous model, which has shown [11] to couple simplicity a satisfactory accuracy in the simulation of the actual HDRB behaviour, has been introduced in the FEM analyses.
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4
Structural analysis of the BIS retrofit solution
The results of the dynamic analyses show the typical, SDOF type behaviour of B.I. structures. From figure 6 it can be derived that the first three modes are the ones characterizing the deformation of the construction acting as a rigid body on the devices. The modal participation factors (table 3) in the longitudinal and transversal directions confirm that the higher modes involving different deformations are practically not effective.
Ti=2.616 s e c
T 3 =2.5 s e c
T2=2.611 s e c
Figure 6. Modal shapes of the BIS retrofit solution LONGITUDINAL PERIOD (sec) M^Mtot(%) 2M,/Mtol(%)
TRANSVERSAL MODE PERIOD (sec) 1 I 2.616
M/M,ot(%) EM,/M 1O1 (%) I 99.102 | 99.102 I
MODE
I
2
|
2.611
I 99.711 I
99.711
Table 3. Modal participation facor: church with BIS retrofit
5
Analysis of results
The results of the analyses presented in this paper are provided in terms of a comparison of the behaviour of the following cases: • monument as is, which means without any retrofit; • monument retrofitted by inserting rigid diaphragms at different levels (the ones given in fig. 4a; • monument retrofitted through Base isolation which dimensions are given in fig.5 and table 2 and which location is given in fig. 4b.
124 25 Vi/Wtot % 0 without retrofit 20
S3 with rigid diapharagms SI BIS
15
10
5 -
Figure 7. Effect of base isolation and rigid floor on shear distribution on longitudinal and transversal elements. Shears nondimensionalized to toal weight of construction.
Due to the limited space not all the results are provided in this paper. Only the distribution of shears among the different transversal and longitudinal elements are given for the three above mentioned cases. The results are given in figures 7 and 8. From figure 7, in which shears, on all longitudinal (LI to L4) and transversal (Tl to T5) elements, are nondimensionalized to the total weight of the church, it can be derived the consistent reduction of forces when BIS is introduced with respect to the "as is" monument. It is also clear that the insertion of rigid diaphragms strongly concentrates the shears on the stiffer elements both in the longitudinal (LI and L3 elements) and transversal (T2 and T5 elements). The same distribution of shears is given in figure 8. In this case the shears on each element are nondimensionalized to the weight of the same element. This representation obviously confirms the same results but gives some interesting information on the efficacy of adopting the retrofit solution of inserting rigid diaphragms. It is in fact demonstrated that, in this particular case, the element in which shears are concentrated: LI and L3 in longitudinal and T2 and T5 in transversal attract shears which respectively are equal to 200 % and 150% of the weight insisting on the element. It can be easily demonstrated that these values of shears cannot be withstood by the element and therefore the retrofit solution is not appropriate unless other interventions are considered together with insertion of rigid floors.
125 Vi/W, %
L1
L2
L2
L3
T1
T2
T3
T4a
T4b
T4c
T4d
T5
Figure 8. Effect of base isolation and rigid floor on shear distribution on longitudinal and transversal elements. Shears nondimensionalized to weight of element.
In conclusion BIS is a very effective means for retrofitting these type of monuments while insertion of rigid diaphragms does not necessarily improve the seismic behavior. 6
Conclusions
The results of analyses of the "as is" church and BIS retrofit solution have shown that: • It is possible to obtain a particularly simple dynamic behaviour of the church through an appropriate design of the BI system. • BIS leads to large reduction of the strength demands in the structural elements, such to avoid tensile stress, and hence cracking, in the masonry elements. These results appear encouraging and suggest that the BIS can be delineated as a particularly promising strategy in the seismic retrofit of historic churches.
7
Acknowledgements
This research has been supported by the CNR Progetto Finalizzato Beni Culturali.
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References 1. Honeck, W. And Walters, M, Use of steel in the seismic retrofit of historic Oakland City Hall. Steel Tips, Structural Steel Educational Council. 1994. 2. Buckle I.G. - "Application of base isolation systems to the seismic retrofit of historical buildings in the United States", in: Final Report on the Int. Workshop on the Use of Rubber-Based Bearings for the Earthquake Protection of Buildings, by James M. Kelly, Report UCB/EERC-95/05, 1995, pp.C5-C18. 3. Mayes R.L., Jones L.R., Kelly T.E. - "The economics of seismic isolation in buildings", Proc. of Int. Workshop on Recent Developments in Base Isolation Techniques for Buildings, Tokyo, Jap., 1992. 4. De Luca A., Mele E. - "The seismic isolation in the retrofit of historic buildings", Proc. of USA-Italy Seminar on the Seismic Restoration of Historic Buildings, Los Angeles, July 22-25, 1996. 5. Wyllie L.A. - "The balance between historic preservation and seismic safety can we achieve it?", in: The Seismic Retrofit of Historic Buildings Conf. Workbook, San Francisco, Cal., 1991, pp.5-1 - 5-9. 6. Elsesser E. et al. "Repair of five historic buildings damaged by the Loma Prieta earthquake.", in: The Seismic Retrofit of Historic Buildings Conf. Workbook, San Francisco, Cal., 1991, pp.4-1 - 4-40. 7. Youssef N. et al. - "Passive control of the Los Angeles City Hall". PVP, Seismic, Shock and Vibration Isolation, ASME, Vol. 319, 1995, pp.241-248. 8. Naaseh S. - "Seismic retrofit of San Francisco City Hall. The role of masonry and concrete", Proc. of the 3rd National Concrete & Masonry Eng. Conf., San Francisco, California, 1995, pp.769-795. 9. Seismic Isolation Project Briefs. Buildings, Dynamic Isolation System Inc., Sept. 1995. 10. Mele E., Modano M., De Luca A. The seismic retrofit of historic masonry buildings through BIS: preliminary analysis for application to church typology. Proc. of MONUMENT '98 Workshop on Seismic Performance of Monuments, Lisbon, Portugal, Nov. 1998. 11. Mele, E., De Luca, A., Ramasco, R. The effect of using different device numerical models on the global nonlinear behaviour of base isolated structures. Proc. of 11th WCEE, Acapulco, Mexico, Paper No. 1541. 1996.
TOWARDS A SAFETY CONCEPT FOR BUILDINGS WITH STRUCTURAL CONTROL
UWE E. DORKA University of Rostock, Dept. Of Civil Engineering, Philip Mtiller Str. 20, 23952 Germany E-mail: uwe. dorka @ bau. uni-rostock. de
Wismar,
For many structural control concepts to work effectively, structures must be modified and the control system then becomes safety critical. The control system modifies the loading on the conventional part of the structure. The reliability of this load modification capability is an important safety issue. Simulation models must be used to assess this load modification capability. They are more refined than design models and must be verified. They are not state of the art knowledge. For a realistic safety assessment, failure of the control system must be taken into account. This adds new failure modes un-common to structural engineers. But not every control failure triggers structural failure. A control system confidence can be defined and its required level may be given in codes. These issues are discussed and illustrated on HYDE-systems, a specific passive control concept for earthquake protection.
1
Introduction
Many structural control concepts have been developed in recent years. They can be classified as passive, active, semi active or hybrid. Because it is a new technology, it is often only used as "add on" to improve performance but not as a safety critical component for the structure. On the other hand, it is more and more evident that, in order to fully utilise the advantages of many structural control concepts, the basic conventional structure should be modified. This often makes the control system safety critical. Take Active Mass Dampers (AMDs) as an example. They are most effective in very slender, light weight structures. But when malfunctioning, they may excite such structures much more easily and may even cause collapse. This example illustrates another important safety issue: The structural control system adds new modes of failure that must be considered, some of them probably yet unknown. This raises the question of how unknown failure modes in new technology can be incorporated in a safety concept. The safety concepts of current building codes like UBC or Eurocodes are based on a separate treatment of loads and local structural resistance. This leads to the concept
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128
of split safety factors. The variability of the loading is taken into account and so is the variability of the material. The uncertainties in modelling the structure is not explicitly included (Fig 1). Since in most cases, only simplified "worst case response" models are used in design (like neglecting the moments due to end restraints of a beam when designing its midspan cross section). It is not common practice (and is of course not necessary for conventional structures) to use a model that can simulate the behaviour of the structure with a certain degree of accuracy. There are no specific provisions in the codes about model accuracy and how it should be verified. Conventional Structure:
Load
^
Structure: Design model
^
Structure: Simulation model
W
Resistance
Structural Control:
Load
P£ mnHifiratinn/
safety check
Resistance
r control svsrem Figure 1.
Design path for conventional structures and comparison to structures with control systems.
In Figure 1 the design path of a structure with control is also given. The control system modifies the loading on the conventional part of structure. A safety critical state is still due to failure of a conventional component (like column or beam), but a malfunction in the control system may cause a loading state triggering such a failure. The reliability of the load modification capability of the control system therefore becomes an important safety issue. In the following, I'd like to discuss the issues raised here in more detail.
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Reliability of Load Modification Capability
The load modification for the structural system is the most important feature of any structural control system. The type of modification depends on the control concept. Let's look at three important passive control concepts for earthquake protection (Fig. 2). Base Isolation (BI) for example filters out a large portion of the loading in the frequency domain. A Tuned Mass Damper (TMD) in contrast draws input energy away from a specific eigenfrequency of the structure and transforms it into its own vibrations. A Hysteretic Device System (Hyde System) on the other hand limits the maximum forces possible to enter a structure and reduces displacements by transforming large amounts of energy into heat. The first two concepts work in the frequency domain and are based on an essentially linear response whereas the Hyde System relies on a strongly non-linear mechanism in the time domain. Each concept works best with different types of structures.
Base Isolation (BI)
Figure 2.
Tuned Mass Damper (TMD)
Hysteretic Device System (Hyde System)
Three important passive structural control concepts: Base Isolation (BI), Tuned Mass Damper (TMD) and Hysteretic Device System (Hyde System).
BI requires a stiff conventional structure with a first eigenfrequency well above the frequency of the BI system. This cancels any vibration in the conventional structure effectively. The remaining low frequency vibration is hardly excited by an
130
earthquake and additionally dampened by the BI system. It becomes a rigid body motion for the conventional structure above the BI system. TMDs require the opposite: Very slender structures with small masses and one very pronounced frequency. The TMD then effectively reduces the response in this frequency, often by an order of magnitude. Typically introduced to reduce service level vibrations, the conventional structure can now be made so slender that an earthquake has hardly any effect. Hyde Systems also require a very soft conventional structure to prevent any damage there. But their primary structure must be stiff-ductile to limit the forces and allow large energy dissipation at small displacements. This reduces overall displacements and keeps the conventional structure in the elastic range. The stiff-ductile primary structure is part of the control system. It contains the Hysteretic Devices (Hydes) [1]. All three systems can easily be transformed into active, semi active or hybrid systems, given there are appropriate devices available. This extension will not change the underlying concept of each control system (and therefore not the basic requirements on the conventional structure) but rather improves its performance or broadens its applicability (like Hybrid Mass Dampers in high rises, where TMDs would become too large). From this brief discussion of various control concepts it is obvious that for each concept to work effectively, a certain basic type of conventional structure should be required and its compliance checked in an application. And to assess the load modification capability in a particular case, structural models are required that can simulate the actual response with an acceptable degree of accuracy. Such "simulation models" must be more refined by nature than the simple "design models" which are the current basis of codes. For a simulation model to be acceptable, it should capture the mean response of the real structure and document its variation. This will allow for an incorporation of the model uncertainty in a probabilistic safety concept. For structures with BI and TMD systems for example, linear models for response evaluations in the frequency domain appear to be appropriate. For structures with Hyde Systems, time domain models should be required because of their strong non-linearities. The structural engineering community is not used to work with simulation models although modern computer technology and software allow their daily use for some time already. In the future, some document written by experts may provide basic modelling guidelines for this purpose. But especially in the introductory phase of a new technology, where we are now with structural control, I believe that each actual
131
model must be checked until experience has provided us with some simple, yet accurate set of modelling guidelines for each control concept. One way to check a simulation model is system identification. For linear models in the frequency domain, inexpensive on site methods are available. For non-linear models, verification is more involved. For Hyde Systems for example, the linear characteristics of the conventional system (SHS) can only be checked without the devices in place. Necessary gaps to allow for the relative displacements imposed at link levels must be checked by inspection. The linear characteristics of the completed structure, including the devices, must also be checked. Having demonstrated the required load modification capability using a verified simulation model that includes a working control system, failure of the control must now be taken into account for a realistic safety assessment. Depending on the type of control system, single or multiple mode failure may cause partial or total collapse of the conventional structure under loading conditions below the design load for a working control system. In active systems, failure is not only "non-performing" but also "over-performing" due to instable control that may occur without any external loading. A control system may have several failure modes not leading to failure of the conventional structure. Therefore, a failed control system does not necessarily constitute a structural failure.
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• • ' f ; r L . . . . : , .~^V, .^rj:
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u Figure 3.
•I'V
' \ \
\
v
Deformation of a 5 storey office building with Hyde system in the ground floor (Hyde system not shown).
132
To assess the safety of a structure with control, the failure modes of the control system and their effect on the structure's response must therefore be evaluated in probabilistic terms. Each mode of failure has a probability of occurrence that is dependent on the quality of the control system. It is a probability that increases with time like the probability of occurrence of the respective design event. This leads to a time dependent safety of the structure. To illustrate this, let's look at a Hyde System under earthquake loading. Fig. 3 shows a typical 5-storey office building with one seismic link [2]. The most important parameter of such a system is the limit force in the link, which is provided by the devices there. It depends on the maximum permissible link displacement. This displacement is determined by the elastic limit of the SHS. It is the true limit state of the building because beyond it, there may be plastic deformation or buckling failure of the slender SHS. STD [mm]
5 00^
Design Point FN = 380 kN
3.88-
|
3.25_ 2.63F[kN]
7.00
200 Figure 4.
1 — J1
1
400
600
—1
800
1
1000
Design curve for link forces based on permissible link displacement STDs for retrofitting large panel buildings with one seismic link [3].
To assess the safety against this permissible link deformation for a nominal link force, link displacement standard deviations (STDs) can be calculated by Monte Carlo Simulation. A simulation with 500 earthquake records generated for the same site specific earthquake will give STDs with a confidence of 95%. Modelling uncertainties can be neglected in this calculations because their effect on the variation of the response is small compared to the variability of the earthquake (and the model should be checked for reasons mentioned above). A safety index P then provides the required distance to the permissible link displacement which may also
133
be a nominal value. This procedure has been suggested in [3] for the design of link forces in retrofitting large panel buildings. Calculating the STDs in this way for several nominal link forces provides a curve with a design point (Fig. 4). In terms of a probabilistic safety concept based on "load" and "resistance" for the conventional part of the building, the "load" is now the maximum link deformation under the earthquake effect (modified by the Hydes in the link) and the "resistance" is the link displacement corresponding to the elastic limit of the SHS. Both can be expressed by their probability density functions, with the overlapping area representing failure (similar representation like in Fig.5). The load PDF is time dependent, but stabilises after a very short time into the earthquake for a well designed Hyde system. It also dominates the failure probability. This justifies the use of nominal values for the permissible link displacement and link forces in this calculations. The approximate assessment of the safety by calculating STDs and using an index |3 is required by the computational effort. New advanced Monte Carlo Methods may allow a more accurate safety estimate in the near future [4]. The next question is: How does the probability of failure of the Hyde System influence the safety of the building? To discuss this in the context of our example, let us assume friction devices have been used. There is a certain probability that those devices may get stuck with time or loose there pre-stressing. Apart from those two extreme conditions, changes in the devices may cause a change in frictional force, up or down. This will increase the variance of the Hyde limit force and raise the probability of failure for the PHS (overload on the PHS). Let us assume that an inspection cycle of x years has been introduced that fully verifies the working condition of the devices at that time (for some friction devices, this cycle may be as long as the expected life of a building). Additionally, the PHS may have one or several serious flaws (including the device connections) that may cause premature failure. From this, we have to look at three basic failure states of the control system: (1) "link stuck", (2) "link loose" and (3) "PHS weak".
"PHS weak" will occur when the strength of the PHS drops below the link limit force. The Hyde system then behaves linearly until it breaks at a certain force level without activating the Hydes. Primary system strength and link limit force can be represented by probability density functions (PDFs) at the end of the last inspection cycle in the lifetime of the building, where the variance in link limit force is largest and the PDF of the resistance of the primary system may have shifted to lower values due to structural deterioration. Because primary systems are conventional structures (rc-walls, steel trusses etc.) their strength PDFs are known. The PDF of
134 the link limit force is a function of the number of devices and their quality. This will lead to a probability that the control system fails (area common to both PDFs, Fig.5) A "stuck link" might damage the primary system severely or break the device(s). The Hyde system will again not be able to activate the Hydes, behave linearly and break, now at a much higher force than in the case of a weak PHS. This case is therefore included in the PDF approach mentioned above: It represents the high end of the link limit force PDF whereas the case of a weak PHS represents the low end of the primary system resistance PDF. i k
/ / /
weak PHS
^
link force PDF
\ I
failure probability
\
/
/
V - " PHS resistance PDF
stuck link
force
Figure 5. Failure probability of control system in a structure with Hyde System.
The failure of the Hyde System will not always cause the collapse of the conventional structure. Above a certain linear resistance of the primary system, the conventional structure will not fail although the control system may. This limit can be calculated using an elastic model for the primary system with brittle link failure. A "loose" link leaves the control system in tact but will have such a low limit force that the displacements in the conventional structure increase to levels that cause collapse there (soft first storey collapse in the building of Fig.3). A non activated control system in this force region is inconsequential to the failure probability of the structure because it will fail in any case. To assess the safety of the building considering the failure of the control system, two failure modes occur now that require two different simulation models: A model for a working control system covers also the case of "link loose" but another model is needed using a linear link restoring force with a brittle failure limit to cover the cases "link stuck" and "PHS weak". With both models the STDs of link displacements can be calculated for various link limit forces. This will provide two curves (Fig. 6). By plotting the PDFs of Fig. 5 on
135 the level of the permissible STD, the confidence that the building will have the required safety level in case of an event (system confidence Cs) can be expressed by C s =l-[p(F 1 ) + p(Fh-F1)] with p(Fi) and p(Fh-F!) as indicated in Fig. 6: pCFh-FO is the joint area of both PDFs bound by F) and Fh. p(Fi) is the area under the link force PDF up to F,. Both areas do not overlap. A required system confidence can now be defined in conjunction with the occurence probability of the relevant event and the required survival probability, suitably at the end of the building's life (end of last inspection cycle). The required value for this confidence may be defined in a code. Depending on the event (earthquake, wind etc.), this value should vary. It may also vary with the use of a building.
link displ. STD
link force
A
PHS resistance PDF
link force Figure 6. Illustration of control system confidence for a Hyde system.
As with other buildings, life cycle cost considerations may enter the definition of a target value for the system confidence. That way the definition can be treated as an opitmization problem. A difficult question to answer in this approach is still the cost attached to failure, e.g. the cost of human life, which has to enter here. In some countries, the cost for a human life is based on the insurance costs (life insurance) and/or production value of an individual. The social status of the occupants of a building enters here which usually drops with the age of a building and its state of maintenance. Such an approach leads to a safety that highly depends on the social status of the occupants and therefore is morally questionable. A minimum
136
acceptable safety level should be required for all buildings. Additional safety and, in the context of structural control, additional system confidence may be based on cost optimization, if the owner so wishes. As can bee seen from Fig.6, the variations of the link forces and the PHS have a significant influence on the confidence level. Thus, an improved quality of the control system (decrease in variations of link force and PHS) will increase the system confidence. With low control quality, only a shift to the right will provide additional confidence but will require larger forces and make the structure more expensive. Because such a shift may not have much of an effect on p(Fh-Fi), a large value here can only be reduced efficiently by improved control quality. This example shows that when dealing with structural control systems, only simulation models capable of estimating the response under various system conditions (including failure modes) will provide an answer to the system's probability of survival. A "system confidence" can be estimated (or required) based on the quality of the control system and the required safety level for the event.
3
Unforeseen Modes of Failure
With new technology, there may be unforeseen modes of failure. Because the main issues of this chapter have been discussed in [5], only the essence of it is reported here. Unforeseen failures may be divided into "forseeable" and "unforeseeable" ones [5]. Two types of unforeseen failures may occur: Those that could have been foreseen, but were not a part of the design team's knowledge and those that were not part of the technical knowledge of the time. Both types of failures can be illustrated by historic examples. An example for a foreseeable failure is the Titanic disaster. With the engineering knowledge of the time, one could have noticed that the section walls were not high enough to prevent the water from overflowing once the ship developed a certain tilt. Those section walls were a new technological achievement but the designer did not develop its full potential because he didn't include such an "exotic" failure mode in his design considerations. It is actually quite possible that, in a first draft, these walls went all the way up but where later reduced in height because of economic or other reasons, with no anticipated failure mode available to prevent that. An example for an unforeseeable failure is the collapse of the Tacoma Narrows bridge. Here, a previously unknown wind induced excitation by vortex shedding combined with the detrimental effect of close bending and torsional frequencies
137
caused a resonance effect that eventually destroyed the bridge. This type of excitation was first observed in this bridge and it took several years of research to fully understand it. A way to improve the safety against foreseeable failures is the incorporation of external expert knowledge. Here, it is important to register the help of experts outside the pertinent field, but with some background knowledge of it. In structural control, two additional engineering disciplines interact with structural engineering: mechanical and electric & electronics engineering. Therefore, in designing a structural control system, one should require engineers from those fields to be part of the design team and have a separate, independent team check all safety aspects of the design and not only its code conformity. Such a "dual" team would certainly want to perform various fault tree analyses which are typically not considered by structural engineers. This will shed some light on possible failure modes that otherwise go undetected. Regular design approaches, like the incorporation of backup systems (e.g. for electronic components) will then allow to provide adequate safety against these failures by minimizing the risk of neglecting "exotic" failure modes. Unforeseeable failures are another matter since they are the result of effects that are not part of the current knowledge. Even the most sophisticated analyses or rigorous testing will not reveal them since those methods are all based on current knowledge. Even scientific study programs are only of limited value in structural engineering: Here, each project has its unique design requirements. There is almost no way to reproduce actual situations in a laboratory or computer model and full scale prototype testing is practically out of the question. Thus, only real life experience will reveal such unforeseeable failures and that experience, with time, will eventually close this safety gap. This experience means possible structural failure and possible loss of life. It is the risk a society takes when new technology is applied. The level of acceptable risk will differ from application to application, even within the same type of technology. An example is transportation technology with different risk perceptions for cars, rail based transportation or air planes. The acceptance of risk in building structures is usually very low: In the perception of society, a building must be safe and sound for a century under all kinds of loading conditions. But there are exceptions. In a catastrophic situation, like hurricanes and earthquakes, nature is perceived as the villain. That's why they are called "natural disasters" and because nature today is perceived as a force that cannot be conquered or cheated upon, the acceptance of risks involved with this type of "loading" is much higher, even in buildings where collapses under such conditions are usually accepted by society. Applications of new technology in
138
buildings to reduce the effects of natural disasters therefore can be based on a much higher risk acceptance level which translates into much lower required safety levels. This perception of society can be used with advantage to foster the application of structural control technology by applying them at first to catastrophic load cases. Let us take passive control systems as example. Hysteretic device systems or base isolation definitely will improve the behavior of most buildings under earthquakes. This has been shown time and again through the application of current knowledge and even recent earthquakes like Northridge [6]. Although some structures with passive systems may collapse in the next strong earthquake, it is safe to say though that most buildings using this type of technology will fare much better than any conventional structure around them and the acceptance of this technology within society will increase with future earthquakes or hurricanes, if enough of these structures are around to prove their advantages in a statistical manner. The unforeseeable failures, which will occur with certainty, will close the inevitable safety gap, but will not distract society from applying such a basically advantageous technology. It is therefore a logical requirement that the safety margin in codes for such systems should not be higher than for conventional systems. That way, many more applications will be possible which is required to produce a statistical mass that, under the next hurricane or earthquake, will produce a representative picture of the advantages of this new technology and provide the necessary experience through unforeseen failures. Restricting its application by excessive requirements on safety might just produce a few sample structures (as is the case today). A failure in one of them then carries an enormous statistical bias which might have a dramatic effect on society's perception of the whole technology and even may prevent its application for a long time to come. Structural control systems are more and more governed by code requirements that are often based on fear of applying new technology, rather than rational safety requirements. Base Isolation has suffered this fate already. It is therefore very important now for the structural control community to commit itself to the development of rational safety concepts for structures with control.
4
Conclusions Drawn for a Safety Concept
Current safety concepts in codes are not sufficient to cover buildings with structural control although the important modes of failure are still defined by the conventional part of the structure. An adequate safety concept must consider the load
139 modification capability of the control system which strongly influences the design of the structure, if an economic advantage is to be gained from the control system. To assess this capability, a simulation model is required for the structure rather than the usually crude design models that are sufficient for the verification of conventional structures. With the current lack of modelling guidelines, those simulation models should be verified for each building to ensure an acceptable model accuracy. The safety of the structure with working control system can then be verified by evaluating the standard deviations of important variables of the conventional part of the structure, like storey drifts or beam end rotations using the verified simulation model. Efficient Monte Carlo methods are available for these calculations, if a time history analysis must be performed (e.g. if strong non-linearities are present). Additionally, the possible failure modes of the control system and their effect on the response of the conventional structure must be assessed. Often, this requires additional simulation models to calculate the response under those control failure conditions. Usually, not all control system failures lead to structural failure. A control system confidence can be defined based on the failure probability of the control system bounded by the limit states for failed and working control, plus the failure probability for a working control system (Fig. 6). The level of required control system confidence may be defined in codes. Life cycle cost optimization may be used in this context but including the cost of human life raises serious moral questions. A minimum acceptable level of confidence should therefore be defined by separate reasoning. Because structural control is a new technology, unforeseen failures may arise. They can be separated in foreseeable and unforeseeable ones. The foreseeable ones can be dealt with by including outside expert knowledge in the design team and have the design approach checked by another team. Unforeseeable failures can only bee detected by experience. This experience may be hindered by excessive requirements on the safety for known failure modes, which does not prevent the unforeseeable ones. Therefore, there is no reason to require a higher level of safety from a structure with control than is required from a conventional structure! Unfortunately, this has not been generally recognised and the wrong approach has been used in modern codes (see Base Isolation). This can only be avoided in the future, if the structural control community commits itself rigorously to the
140
development of a comprehensive safety concept for these exciting and valuable structures.
References 1. 2.
3.
4.
5.
6.
U.E. Dorka; Hysteretic device systems for earthquake protection of buildings. 5'h US nat. conf. on earthq. eng. (5th NCEE), Chicago, 111, 1994. U.E. Dorka, V. Bayer; Distribution of seismic links in Hysteretic Device Systems. 12th World conf. on earthq. eng., (12th WCEE) Auckland, New Zealand, 2000. U.E. Dorka, A. Ji, E. Flygare; A hysteretic device system for earthquake retrofit of large panel buildings. 11th European conf. on earthq. eng. (11 th ECEE), Paris, France, 1998. V. Bayer; Ch. Bucher, U.E. Dorka; First passage reliability of bridges by spectral importance sampling. Proceedings EURODYN'99, Prag, Tchech Rep., 1999. U.E. Dorka; New technologies and codes: Or how to deal with the "unforeseen". 8' int. conf. on appl. statistics and probability. (ICASP8), Sydney, Australia, 1999. J.F. Hall, W.T. Holmes, P. Somers (Ed.); Northridge earthquake reconnaissance report, Vol. 1&2, Earthquake Spectra^ Vol. 11, Suppl. C, April 1995 & January 1996.
SEMI-ACTIVE CONTROL OF 3-D LINEAR AND HYSTERETIC STRUCTURES FOR SEISMIC APPLICATIONS
SAMIEL-BORGI Associate Professor, Ecole Polytechnique de Tunisie, BP 743, La Marsa 2078, Tunisia E-mail: sami. elborgi@ept. rnu. tn CHOKRIZAMMALI Graduate Student, Ecole Polytechnique de Tunisie, BP 743, La Marsa 2078, Tunisia E-mail:
[email protected] PANOS TSOPELAS Assistant Professor, The Catholic University of America, Washington DC 20064, USA E-mail:
[email protected] A number of studies have been conducted so far on active or hybrid control of hysteretic structures using various linear and nonlinear control algorithms. However, most research work on the use of semi-active control systems has been limited to linear structures modeled as 1-D shear-type representation (one degree of freedom per floor). The structures used in this study are modeled with a 3-D shear type representation (three degrees of freedom per floor: two translations and one torsional rotation), and can behave either in a linear elastic or in a hysteretic manner. The purpose of this paper is to examine the effectiveness of semi-active variable viscous dampers ( W ) in reducing the seismic response of such structures. The control algorithms considered is one based on the Sliding Mode Control (SMC) algorithm. The analysis results of a linear and a hysteretic, single-story 3-D structures indicate that W dampers can be effective in reducing the displacement and acceleration responses of both linear and hysteretic structures.
1
Introduction
Semi-active control devices combine the features of active and passive control to reduce the response of structures to various dynamic loadings. The main purpose of this paper is to examine the effectiveness of semi-active variable viscous ( W ) dampers in reducing the response of 3-D structures subjected to relatively strong earthquakes. Several investigators have developed algorithms for selecting the appropriate damping parameters of these devices during the structure's response to an excitation. The algorithms included a clipped optimal control algorithm [7], a bang-bang algorithm [3], a Linear Quadratic Regulator (LQR) algorithm [6,7], a Sliding Mode Control (SMC) algorithm [9,11], a generalized LQR algorithm with a penalty on the acceleration response [8], a displacement-acceleration domain algorithm [8] and a fuzzy logic controller [6].
141
142
Under relatively strong earthquake excitations, structural members might experience yielding and the response will become non-linear hysteretic. To date, numerous studies have been conducted on active or hybrid (active/passive) control of hysteretic structures modeled with a 1-D shear-type representation using a wide range of linear and nonlinear control algorithms [10,11]. However, most research work on the use of semi-active control systems has been limited to the same type of structural model accounting only for linear elastic behavior [8,9]. On the other hand, few researchers considered the control of 3-D structures with linear behavior only [1,4,5] except for [6] who studied base-isolated structures with a hysteretic behavior for the base and a linear elastic behavior for the super-structure. In this paper, the structure is modeled with a 3-D shear-type representation and can behave either in a linear elastic or in a hysteretic manner. The algorithm for selecting the damping properties of W dampers presented here in is the SMC algorithm, which is formulated based on full-state feedback. The equations of motion are formulated in the state-space and in the drift (inter-story) coordinate system. A linear and a hysteretic, single-story structure subjected to the El Centro earthquake applied along the F-direction are analyzed using the SMC algorithm. 2
Equations of Motion of Controlled Structure
Consider an n-story structure modeled as a 3-D shear-type representation equipped with m W Dampers, placed each in the X or y-direction, and subjected to the two horizontal components of an earthquake ground excitation ag = {x ,yg). Each story is described by three degrees of freedom as shown in Figure 1: two horizontal translations and one torsional rotation. This structural model was adopted from [6]. If the behavior of the structure is hysteretic, its motion is described by the following system of differential equations defined in the physical coordinate system: MX + CX + KelX + Kin V = DU + Eag
(1)
X = (*[ y\ 9] ••• x„ y„ 6„f a 3n-vector with x}, yj and 0, being the relative interstory displacement of the j * story in the X, Y and 0-direction, U is an m-vector representing the control forces generated by the W dampers, M and C are respectively the mass and damping matrices of size (3n x 3n), D is a (3n x m) matrix defining the locations of the control forces, Ket and K,„ are, respectively, the elastic and the inelastic stiffness matrices of size (3n x 3n), E is a (3n x 2) mass matrix, representing the influence of the earthquake excitation, whose elements are zero except: £(3/ +1,1) = -MXM and £(3/ + 2,2) = -MyM for i=0,1,... ,n-1. All elements of M, C, Kei and K,„ are zero except:M(i,j) = M, for j
o a OP 05
Figure 5: the effect of aQ on the noise reduction
Ambient vibration Daia Random Decrement Vibration Signature
c
I'll o vibialiimie.-.i'on'.e r> Ibraham Time Domain Method j Mode shapes-damping ratios and frequency III' Frequency Domain response , Filtering the low frequency mode shapes;
Figure 6: Outline of the 1st method
IFFT
10th mode
19th mode
; - -1 A7\H/: Distance from main tower Figure 7: Mode shapes
222
Figure 8: The comparison of the natural frequencies In this paper, time domain technique is used to identify the dynamic properties of Hakucho Bridge. This technique can be used without any assumptions for the exciting force, the structure damping and/or the space between the natural frequencies.
2.3.1 Ibrahim Time Domain Method (ITD) Free response with multi degree of freedom system that obtained in the previous step is supposed to be a liner sum of free vibration response of each mode. Ibrahim time domain method is used to determine the natural frequencies, damping ratios, and mode shapes. The system is assumed to be described by the following equation during its free response:
[MP}+[CM+M*}={O} where [M], [C], and [K] are n x n matrices, while X,X, dimensional vectors.
(4) and .Xare n-
The solution of this equation is: 2/V
.
*(0 = 2 > p ( K > r +korj\-£)t){\i/r}
(5)
r=l
Equation (4) can be expressed in terms of matrix form by discretizing the response x at time tAj
[X] = m [ A ]
= 1,2,..., L) as,
(6)
223
Where[X] = [JC, (t})],
[V] = [ ^ ] ,and
[A] = [exp((-£>, + iQ), <Jl-tf)tj)]. The dimension of each matrix is, [X]: n x L, [\|/]: n x 2N, and [A]: 2N x L, where n is the number of measurement points, and L is the total number of time steps. The response with delay time of At is expressed as,
x(t) = x(t + At)
(7)
Which can be discretized to,
[X] = PP][A]
(8)
An n x n square matrix [A] to transform from the original time history to the lagged time history is defined as,
[Aim=m
(9)
Using this relationship, the modal properties can be obtained by solving the following eigenvalue problem:
[A]{Wt} = exp((-£>, + ia), VW?)AO{^} (10) The matrix [A] is derived by pre-multiplying both sides of equation (6) by [A] to obtain,
[A][X] = [A]PF][A]
(11)
Which yields to, [A][X] = [X]
(12)
Post-multiplying equation (11) by pseudo-inverse of [X], [A] can be obtained as, [A] = [ X ] [ X f ( [ X ] [ X ] r r 1
(13)
Because all the modes are identified simultaneously by solving the eigenvalue problem of equation (10), this method can easily be applied for multi mode structures with closely spaced natural frequencies, such as suspension bridges.
224
2.3.2 Identify the Higher frequencies Modes
g-
.jtz^T*'
f-0
0.5
2
1.5
2.5
3
Acceleration response rms Figure 9: Bridge response and wind speed dependency
a) The 1st and 2na mode ,„..,,.,..„.,
! • • • £
3.18
-__ *
"
m.d.-7
*— ^___ •
OS
* *^ _ !,:
1
^ - : -—__
'• 2
1.5
•
I s
*
^~* * 0 5
1
IS
i
2 s
b) The 17th and the 18th mode Figure 10: Natural frequency and amplitude dependency
225
a) The 1st and 2na mode
b) The 17th and the 18th mode Figure 11: Damping ratio and amplitude dependency
226
1
1
i:
i
i
1
Lx-^*!
^p
-i - - r
S1
- M£d=-' -
i
i
1
i
i
i
i
1
i
i
i
i i i
i
i
i 1
—
i
i i
mod«2
i
0.2
1
.
0
1 r
•
-0.2
04
J
x ^ 1
-1 1
1
:-i--
I
— i —
-
P5
I
i
i i- -
1
1 1 1 1
—
s.
1 1 1 1
*r
.. i. ...
i
i
i U
1
&
i "
1"
- -1
(
r
k
i
i 1 I. -1- - - 1 r
i
I
Distance from main tower (m) a) The 1st and 2nd mode
Distance from main tower (m) b)The 15th and the 16th mode Figure 12: Mode shape and amplitude dependency
227
By direct application of ITD only the first few modes of low frequency can be identified with high accuracy, however the higher modes can only be identified with less accuracy [81. To use the dynamic properties in health monitoring of structures, the higher modal parameters should be identified with high accuracy. In this study the following two different methods are used to identify the high frequency modes accurately: 1. By filtering the low frequency components. 2. By increasing the number of measurements (Time Shift Method). Filtering the Low Frequency Components Outline of the scheme used in this method is shown in (figure 6). The high frequency mode shapes are identified beginning from the ambient vibration data. The following procedure is used in this step: i) Transform the time domain response to frequency domain response by using the Fast Fourier Transform (FFT). ii) Filtering the first part, which contain some of low frequency modes by using high pass filter. iii) Return the response again to time domain by using the Inverse of Fast Fourier Transform (IFFT). iv) Identify the high frequency mode shapes by using Ibrahim time domain method. Repeat these steps until identify all the higher mode shapes. Time Shift Method By using equation (5), which will be rewritten in equation (14), at any instant of time (t), 2N conjugate modes are identified. But only the first few modes can be identified with high accuracy.
'*i(0 x2(t)
2N
x
lr
T
2r
r=l
xN (t)
.
•exp((-£> r +>/l-C>r)0
Nr .
•(14) After time shift x, the same 2N modes are identified as in equation (15)
228
f^l
X^t + Tj 2N
X2(t + Tl)
* 2 , •
=
!
•
xN(t + tx)
•exp((-Crcor+jl-Ccor)
^Nr
(t + Tj)
.(15)
Equation (15) can be rewritten as in equation (16):
[*.r] X2(t + T1)
2N
-2-
®2r
• exp((-£>r + 4l~Co)r )t)
r=\
XN(t + Tl)
®Nr
(16) Then the number of measurement points increased from N To 2N and the number of modes also increased from 2N to AN. By making the same time shift many times and combining all equation in one equation, the number of modes extended from 2N to 2L as shown in equation (17): ^
T,2r
x2(t)
2N
xN(t) xx (t +
T)
Y.Nr V.(N+l)r
exp((-£> r +
Y,(,N+2)r X2{t + T)
*¥Lr XM(t + T)
4l-Cr= £ % >
0.00 -0.05 -0.10 -0.15 "°' 2 8.0
0.5
1.0
1.5
2.0
Time [sec]
Fig. 16. Speckle noise Identification Method of Mode Shape without Knowledge of Input Excitation The measured vibration results using LDV can not be measured all points at the same time. Furthermore ambient vibration that is used as input is random vibration, which possesses non-stationary, so usually mode shapes can not be identified. Then following averaging methods are carried out. At first, the effect of non-stationary in measured ambient vibration is dispersed by getting shorter of the measurement time for one point and sweeping the area of interested. Next, temporary averaging that random components is stacked the time histories recorded repeatedly is carried out. In these way, stationary vibration components are extracted from ambient vibration without knowledge of input excitation, and then mode shape is identified by counting peak amplitudes of spectrum. Experimental Verification The identification of mode shape for a steel plate (385x300x2 [mm]), which was fixed at one side, was demonstrated experimentally. With respect to measurement condition, measurement points were arranged as 20 points along vertical direction and 15 points along horizontal direction, and total 300 points were set. Measurement time was 2048 [sec] per point considering the step size of frequency and the length of measurement time.
234
Sampling frequency was decided as 2000[Hz] from objective modal frequency. Then, the measured vibration results were recorded into the a personal computer through an AD converter. On the other hand, the change of measurement point was controlled by the personal computer through a DA converter. In this measurement, the maximum irradiation angle was 4 degrees. Based on this measurement condition, measured velocity amplitude spectrum was averaged by 300 times repetitive measurement. The 5 times standard deviation per each measured vibration result was employed to judge speckle noise. Fig. 17 shows the comparison of averaged velocity amplitude spectrums between 1 and 300 times repetitive measurement. From this figure, noise level was suppressed by the averaging process, and the peak of high frequency vibration component was stuck out clearly. The identified 6th and 14th mode shapes, which were changed in repetitive number [10,100,300], and theoretical analysis results of the plate by RayleighRitz method [8] are shown in Fig. 18. The mode shapes were normalized that maximum value became 1, and the absolute values were plotted. It is clear that the identified mode shapes are closely to the theoretical results, as the number of repetitive measurement is increased. Therefore it can be considered that nonstationary and random component in measured ambient vibration are eliminated by 300 times repetitive measurement. Table.2 shows the comparison between identified modal frequencies and theoretical ones. From this table, 10 mode shapes until 16th mode were identified in this vibration measurement. Moreover, in Fig. 17, many peaks except for vibration components are stuck out. These are caused by vibration component of LDV, electric noise and so on. For example, these peaks are extracted and plotted in Fig. 19. This shape is not obviously different from mode shape of the plate. Therefore mode shapes and the other noise components can be distinguished visually using this proposed method.
235
io3r g io2 o &
i
# io1 "I 10°
T« im«"HWFm»rinw 'IV'HPIW1! "Hf^nnp ••
>
M io"'0
n
r» j
200
o
P*
™•" I P '
400
'"• r
"T ^ ^ • T
600
800
'•f
T
1000
Frequency [Hz] (a)
103r
B io o &
^
io
,
I 10° >
I
Ill 1 UiuJ iVliuuji.
.§ io"10 o
200
400
111 1
600
800
1000
Frequency [Hz]
(b) Fig. 17. Comparison of averaged velocity amplitude spectrum: Repetitive number (a) l;(b)300
236
Fig. 18. Comparison of identified mode shapes and theoretical ones: (a) Repetitive number: 1; (b) Repetitive number: 300; (c) Theoretical mode shapes (6th mode; above, 14th mode; bellow)
Mode order 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16
Table 2. Identified Modal Frequencies Theoretical Value [Hz] Identified Value [Hz] 10.74 12.07 34.18 76.01 66.41 121.0 112.8 137.2 135.0 192.4 212.7 234.5 233.9 260.9 336.7 391.6 384.8 371.1 417.7 432.5 465.6 580.1 599.5 606.0 666.5 686.4
237
Fig. 19. Vibration shape due to noise
3.3 Damage Detection on the Change of Mode Identification Method of Change of Mass and Stiffness Caused by Damage The effect of structural damage comes out as the change of mass and stiffness of structures. So the quantitative knowledge of mass and stiffness decrease is needed to evaluate damage degradation. Hence, the damage detection method based on the change of mode shape before/after damage is suggested. The merit of this method is that information of material characteristics, geometrical shape and boundary condition is not needed if modal frequencies, mode shapes before/after damage and mass density are obtained. That is, the damage detection can be carried out by identified modal frequencies and mode shapes before/after damage using the spatial vibration measurement method, if mass density is obtained. The theoretical procedure is shown as follow. At first, it is considered that negative mass and stiffness are added to the objective structure by damage. Considering that the structure before damage, additional mass and stiffness are separated, total kinetic energy : T and potential energy : Vare shown in the following equation [9].
Z
Z
i=l
£ ,=1
± 7=1
7=1
£ ;=i
*• y=i
In these equations, i is mode order, N is total mode number. Mt, Kt, CO,- and qt show modal mass, modal stiffness, modal frequency and modal displacement
238
ofrthmode, respectively. On the other side, j is the number of additional mass and stiffness, J is total number of those. mp kt and «j show jth additional mass, stiffness and the displacement at point j respectively. Further q,it means first derivative of q and u with respect to time. However, the amount of J numbers of restrict conditions, which are shown as following equation, have to be considered at each additional location j of mass and stiffness. jr<J>,.OX.-",. = 0
(.7=1,
,J) (20)
1=1
Hence, /(/) is zth mode shape before damage at point / Considering the restrict conditions of Equation(20), Lagrange function is given by next equation.
L = r-v + X^|s*/0"to,-«,}
(2D
Aj is Lagrange's unknown parameter. By Lagrange's motion equation,
ML
=0
(22)
dt 34, is given, then the following two equations were obtained by expanding Equation(22) with respect to Uj, the same as gt. M& +a>i2Miqi - 2 ^ , O " ) = 0 mjuj+kJuj+
l
XJ = 0
Hence, qi='qie"a, Equations(22),(23),(24),
uJ=uJe"a,
(24) Aj =Aje"a are
-0}2M^ +0)i2Miqi - X X ^ ( ; ) = 0 -co2mjZj + kjZj + Ij = 0 E^,O')^,-Z7=0
(23) substituted
into
(25) (26) (27)
i=i
were given. Next equations are derived as z is eliminated by solving equations given by Equations(25),(26),(27).
239
+ 7=1
(28)
i=l
•or % + X ^ E ^ ^ K O " ) =0 7 =1
i=l
Therefore, the calculation of modal frequencies and mode shapes of the structure after damage is come back solving eigen problem given by following equation [10]. A-COd
B v = {0}
(29)
A and B are matrices(/VxAO, the components for each matrix are shown by following Equations(30), (31). And CO is equal to CO, v is the N order vector}^ •••g„}' . mn
tnn
m
m
^^ j=i
j
m ^ •/ /
n
^J'
(30)
Bmn-SmnMm+^m^mU)^nU) ;'=i
(m = l,
,N, n = 1,
,N)
(31) In the above equations, S is the Dirac's delta function. Therefore, ith modal frequency after damage is given by CO ,-, mode shapes O, after damage is obtained by
(32) So, unknown parameters, which were needed to calculate mode shape after damage, are only additional mass, stiffness and location, if the distributed mass of structures and mode shapes before damage are obtained. rrij and kj are identified by minimizing the following evaluation function(33), that is the difference between the mode shape : W and modal frequency : E7 after damage measured by spatial vibration measurement and ones obtained analytically is minimized.
£=S
£[¥,.(*)-1. Using a TLCD with a total liquid length L of 0.25m produces a response spectrum where the dynamic amplification factor has a local maximum of 18.3% of the original structure without TLCD. This is a local maximum recorded at a frequency ratio of 0.974. An additional local maximum is seen to exist at a frequency ratio of 1.076 equal to 7% of the original value. Since the two local maximums are different in value and the larger one lies at a frequency ratio less then one, the natural frequency of the TLCD must be lowered (see Figure 6). Therefore, by increasing the total liquid length L to 0.27m an improvement is expected. The first local maximum occurs now at a frequency ratio of 0.949 equal to 13.3% of the original value. The second local maximum equal to only 10.3% lies at a frequency ratio of 1.058. The liquid length is now increased to 0.275m and the first local maximum reduces further to 12.2%, while the second maximum increases to 11.3%, always of the original undamped response. Thus, the optimal value for L is very close to 0.275m. The last experiment is performed using L equal to 0.28m.
Excitation Frequency [Hz]
Figure 5: Frequency Sweep - Shear Frame without TLCD
289 As expected, the first local maximum is now lower then the second one, 10.35% compared to 12.7%. Using an interpolation procedure of these results, the optimum liquid length L is found to be equal to 0.276m, corresponding to a natural frequency of the TLCD equal to 1.34 Hz. Thus, the frequency ratio to provide optimum vibration mitigation is equal to 0.979. This result corresponds with the findings of 0.98 in Hitchcock et al. [8].
0.70
0,80
0.90
1.00
1.10
1,20
1,30
1.40
1,30
1,40
F r e q u e n c y R a t i o : Excitation F r e q u e n c y / E i g e n f r e q u e n c y
Figure 6: Frequency Sweeps - Variation of the Liquid Column Length L - Model 1A
0,70
0,80
0,90 1,00 1,10 1,20 F r e q u e n c y Ratio: Excitation F r e q u e n c y / E i g e n f r e q u e n c y
Figure 7: Frequency Sweeps Models 1A-1B-1C
290
0,70
0.80
0,90
1.00
1.10
1,20
1.30
1.40
Frequency Ratio: Excitation Frequency / Eigenfrequency
Figure 8: Frequency Sweeps Models 1C-1D
The final investigation focused on determining the best value for B using a constant liquid length L of 0.275m. Model 1A, IB, IC and ID are used for frequency sweeps and the results shown in Figure 7 and Figure 8. It is interesting to note that the two local maximums are now approximately equal for Model 1A, IB and IC. In other words, the optimal liquid length L does not change for different geometry, see Figure 7. It is to be further noted, that increasing the length of the horizontal column does increase the damping characteristics of the TLCD. Thus, Model IC with a total liquid length L of 0.275m does provide the best damping from all models analyzed. The maximum value of the dynamic amplification factor reduces to only 9.3% of the original response (100%). Comparing the results of Model ID with the results from the previous three models, it shows that increasing the horizontal column length to 0.23m does not improve the response any further. For a total column length of 0.275m, only 0.0225m of fluid are in each vertical column when the TLCD is at rest. During motion, the horizontal column is not continuously filled with water, introducing large additional damping due to sloshing in the curved part of the tube, shifting the eigenfrequency of the TLCD. Thus the coupling between the primary and secondary structure weakens due to the shifted frequency ratio and the effectiveness of the TLCD is reduced. Even though the damping of Model ID is increased as determined during the free vibration tests, for the frequency sweep in Figure 8, the calibration of the TLCD is found to be not optimum.
291 4
Conclusions
Results from free vibration tests and unidirectional frequency sweeps are summarized in this paper with the objective to determine the optimal total liquid length L and to evaluate the influence of the horizontal column length B on the damping characteristics of a small scale TLCD. It is found that the results are best when a frequency ratio of natural frequency of the TLCD to the eigenfrequency of the SDOF structure is equal to 0.979. Further, the horizontal column length is important and in general the damping does increase with an increase of B. However, if the geometry of the TLCD is such that during motion the horizontal column is not filled with liquid at all times, the coupling between the two structures reduces the energy transfer between the primary and secondary system and effectiveness of the TLCD is reduced. References 1.
Balendra T., Wang CM. and Cheong H.F., Effectiveness of tuned liquid column dampers for vibration control of towers. Engineering Structures, 17(9), pp.668-675. 2. Blevins R.D., Applied fluid dynamics handbook. Van NostrandReinhold, 1984. 3. Chang C.C. and Hsu C.T., Control performance of liquid column vibration absorbers. Engineering Structures. 1998, 20(7), 580-6. 4. Chang C.C, Hsu CT. and Swei S.S., Control of buildings using single and multiple tuned liquid column dampers. Structural Engineering and Mechanics. 1998, 6(1), pp.77-94. 5. Fried E. and Idelchik I., Flow Resitance: A design guide for engineers. Hemisphere Publishing Co. 1989. 6. Haroun M.A., Pires J.A. and Won A.Y.J., Suppression if environmentallyinduced vibration in tall buildings by hybrid liquid column dampers. The Structural Design of Tall Buildings 1996, 5, pp.45-54. 7. Hitchcock P.A., Kwok K.C.S., Watkins R.D. and Samali B., Characteristics of liquid column vibration absorbers (LVCA) - I. Engineering Structures, Elsevier Science, Vol. 19, No. 2, 1997,pp.l26-134. 8. Hitchcock P.A., Kwok K.C.S., Watkins R.D. and Samali B., Characteristics of liquid column vibration absorbers (LVCA) - II. Engineering Structures, Elsevier Science, Vol. 19, No. 2, 1997, pp.135-144. 9. Hruska A., Elastic Shear Frames with Tuned Liquid Column Dampers, M.Sc. Thesis, Technical University of Vienna, 2000. 10. Iwan W.D. and Yang I.M., Application of statistical linearization techniques to nonlinear multi-degree of freedom systems. J. Applied Meek, 39, pp.545-550, 1972.
292 11. Kwok K.C.S., Xu Y.L. and Samali B., Control of wind-induced vibrations of tall structures by optimized tuned liquid column dampers. Cheung Y.K., Lee J.H.W. and Leung A.Y.T. (eds), Computational Mechanics, Balkema, Rotterdam, 1991, pp.249-254. 12. Sadek F., Mohraz B and Lew H.S., Single and multiple-tuned liquid column dampers for seismic applications. Earthquake Engineering and Structural Dyanamics, 27, 1998, pp. 439-463. 13. Sakai F., Takaeda S. And Tamaki T., Tuned liquid column dampers-new type device for suppression of building vibrations. Proc. Int. Conf. on Highrise Buildings, Vol 2 , Nanjing, China, 1989, pp. 926-931. 14. Saoka Y., Tamaki T., Sakai F. and Takaeda S., A proposal for suppression of structural vibrations by tuned liquid column dampers. Proc. 43rd Ann. Conf., JSCE, 1988 (in Japanese). 15. Swaroop K.Y., Kareem, A. and Kantor, C.J., Semi-Active Control Strategies for Tuned Liquid Column Dampers to Reduce Wind and Seismic Response of Structures. 2nd World Conference on Structural Control, Kyoto 1998 16. 2nd World Conference on Structural Control, Kyoto 1998 17. Tuned Liquid Column Dampers to Reduce Wind and Seismic Response of Structures" 18. 2nd World Conference on Structural Control, Kyoto 1998 19. Won A.YJ., Tuned Liquid Column Dampers and Hybrid Liquid Column Dampers to suppress earthquake-induced motions in flexible structures. Ph.D. Dissertation, University of California Irvine, 1994. 20. Xu Y.L., Samali B. and Kwok K.C.S., Control of along-wind response of structures by mass and liquid dampers. J. Eng. Mech. ASCE. 1992, 118, pp.2039. 21. Yalla K.S., Kareem A. and Kantor D.J., Semi-active control strategies for tuned liquid column dampers to reduce wind and seismic response of structures. 2" World Conference on Structural Control. Kyoto 1998.
DEVELOPMENT OF A TESTING SYSTEM FOR DYNAMIC CHARACTERIZATION OF ENERGY DISSIPATING DEVICES
HIROKAZU IEMURA, A. IGARASHI AND H. TANAKA Dept. Civil Engineering Systems, Kyoto University, Sakyo-ku, Kyoto 606-8501, Japan E-mail:
[email protected],
[email protected] The objective of this research is to develop a system which allows an economical and accurate loading of dampers and energy dissipating devices to characterize the dynamic properties and performance of those devices under the loading condition when it is installed to structures subjected to strong earthquakes. The concept of the test is an elastically supported main mass system and a mass-driver device (shaker) installed on the main mass system. This system forms a SDOF system, which is dynamically excited with the massdriver system either with sinusoidal or seismic input. The proposed method is shown to be an effective test procedure and it has great advantages over the conventional testing method by verification test using an oil damper specimen to demonstrate the capability of the test system.
1
Introduction
Recent findings and investigation on the damage of civil engineering structures due to strong earthquakes revealed the essential need of the concept of seismic isolation and structural control, especially with added damping mechanisms by means of energy dissipating devices in order to achieve the acceptable earthquake performance of the structure. For this reason, various dampers and energy dissipating devices have been developed for the actual application to buildings and large-scale structures. However, verification of such devices with conventional loading equipments, typically with hydraulic systems, requires a high level of capacities to the hydraulic system with significantly large strokes and velocities. The objective of this research is to develop a system which allows an economical and accurate loading of dampers and energy dissipating devices to characterize the dynamic properties and performance of those devices under the loading condition when it is installed to structures subjected to strong earthquakes. 2
Method and Experimental System
The concept of the testing system consists of a RC floor system supported by roller bearings and an elastomeric bearing to provide elastic restoring force and a massdriver device (shaker) installed on the floor system. Mass of the RC floor combined
293
294 with the elastic restoring force-forms a SDOF system, which is dynamically excited with the mass-driver system.
Shaker
Figure 1. Concept of the test system
Figure 2. Side view of the test system
The shaker, which consists of an auxiliary mass and a mass driving mechanism (an electric motor connected to the mass with a ball screw shaft), is controlled so that the dynamic response to the designated ground motion is induced at the floor system. Inertia force generated by the vibratory motion of the floor is' transmitted to the damper specimen. There are several advantages in this, method; the loading time history for the energy dissipating device reflects the boundary condition (i.e. displacement, velocity and acceleration) to the device when it is installed in the main
295 structure under earthquake excitation, including the dynamic interaction between the device and the structure in the real time scale.
*.& ^m<m^ y Figure 3. Mass Driver System for the excitation of the test system
Tables 1 and 2 show the specification of the main floor mass system, and that of the mass driver system, respectively. Table 1. Specification of the test system.
Main Mass Weight Dimensions
26.9tf 7.0m X 4.0m X 3.0m
Table 2. Specification of the mass driver system
Auxiliary Mass Weight Maximum Stroke Maximum Driving Force
2.0tf
"
1
±50cm 15.7kN
Loading capacity of this system is theoretically evaluated considering stroke limit of the mass driver system as well as fundamental design parameters such as the mass of the mass driver and the main floor system, maximum driving force of the mass driver, and the equivalent damping of the device.
296
"0
0.477
1 0832
frequency[Hz]
Figure 4. Loading capacity of the test system
3
Verification Test
Theoretical evaluation shows that effective loading tests can be performed with die proposed test concept. In order to check the capacity of the testing system used with earthquake record inputs, numerical simulation of the test process, including the stroke limiting algorithm and driving force restriction is conducted, and it is shown that the loading condition is acceptably accurate.
3.1
Test Specimen
The energy dissipation device model used as the test specimen is an oil damper intended for the use of seismic response reduction of building structure, shown in Figure 5. The damper is designed to generate nonlinear viscous damping force in accordance with the characteristics shown in Figure 6.
297
Figure 5. Oil clamper specimen
Velocity Peine]
Figure 6. Design characteristics of the oil damper specimen Table 3. Specification of the oil damper
Damping Coefficients 1 Maximum Stroke 3.2
Ci=100, C2=29.4 [kN/(m/sec)]
±5cm
Sinusoidal Loading Test
Figures 7 and 8 show the results of sinusoidal loading tests, in which the damper is subjected to sinusoidal loadings with various frequencies and amplitudes. For the test cases in which the velocity amplitude of the loading is +/- 3cm/sec (Figure 7), the behavior of the damper is expected to be linear, and for the case +/- 6 cm/sec (Figure 8), the yielding takes place in the velocity-damping force relationship. The test results indicate that the energy dissipation performance is as expected, except that hysteresis loops are observed in the velocity-damping force relationships, presumably caused by associated mechanical phenomena such as friction and backlash as well as damper stiffness.
298 Damping force-Disp.(0.5Hz lOgal)
*
0
Damping force-Vel.(0.5Hz 10gal)
£
0
-5 -2 Damping fon#tfep'!fe.5Hz 60gal>
£
0
-5
^ 0
-0.2
0
0.2
-5
0
Disp. [cm]
Vel.Ckine]
Figure 7. Sinusoidal loading test results for linear region (top: 0.5Hz, lOgal; bottom: 3.5Hz, 55gal) Damping force-Disp
Damping force-Vel.
f. o
-10
0 DampVreOMP&l-Vel.
•2 0
0
2 - 1 0
0
disp. [cm] vel.fcine] Figure 8. Sinusoidal loading test results for nonlinear region (top: 0.57Hz, 15gal; bottom: 1.5Hz, 50gal)
299
3.3
Seismic Loading Test
Seismic loading tests are conducted by controlling the mass driver so that the response of the main mass system simulates that of the structural system excited by earthquake ground motion. Figure 9 shows a typical set of seismic loading test results, in which JMA Kobe record NS component (scaled to 50gal maximum acceleration) from the 1995 Hyogo-ken Nanbu earthquake is used as the input. In spite of slight difference between the design and the actual damper response, especially in the force - velocity relationship, the response reduction effect of the device is well predicted.
Damping force-Disp.
Damping force-Velocity
Exp Sim
•6 0
•2
0
s 1,me History o Displacement
disD.fcml
' !
A ?! 1 \ IK
IK
vellkinel
S^w-~^/^-4^~~/>**s^^
i "
i
i
Time History of Vel.
Figure 9. Seismic input loading test results (Kobe record NS, 50 gal max. ace.)
300 4
Conclusions
The proposed method is shown to be an effective test procedure and it has great advantages over the conventional testing methods. One of the clear advantages is that the method is economical and effective in testing real size energy disspating devices. Also, the loading time history for the energy dissipating device reflects the boundary condition (i.e. displacement, velocity and acceleration) to the device when it is installed in the main structure under earthquake excitation, including the dynamic interaction between the device and the structure in the real time scale. Construction of an actual example of the proposed test system has been completed. This system has been accepted as the Japanese patent. Experiments of different types of dampers with the constructed system are now being carried out. Following the passive devices, semi-active and active dampers will be tested.
5
Acknowledgements
The research described in this paper is supported by Japan Society for the Promotion of Science. References 1. 2.
Soong, T. T. and Dargush, G. F., Passive Energy Dissipation Systems in Structural Engineering (Wiley, Chichester, 1997). lemura H., Igarashi A. and Toyooka, A., Study on simulation of earthquake response of structure using excitor, J. Structural Engineering 45A, (Japan Society of Civil Engineers, 1999), pp. 719-726 (in Japanese).
PERFORMANCE OF VARIABLE-DAMPING SYSTEMS: THEORETICAL ANALYSIS AND SIMULATION J.A. INAUDI National University of Cordoba, Argentina Email:
[email protected] Abstract The dynamic performance of structures with variable-damping actuators subjected to broad-band excitation is analyzed herein through numerical simulation of simple models. A comparison of the response of structures with passive, active and semiactive controllers is developed to evaluate the efficacy of semi-active damping devices for vibration reduction. It is demonstrated that semi-active dampers can achieve better performance than passive dampers and comparable to fully active systems in the case of lightly-damped frame structures. Using random vibration concepts, the relative efficiency of the realization of clipped linear optimal controller using semi-active dampers is evaluated in conventional structures, structures with tuned mass dampers and in interacting structures. New algorithms for variable friction and electro-rheological (ER) actuators are presented aiming at deformation reduction and acceleration reduction. The use of semi-active ER dampers as enhanced passive dissipaters that require deformation feedback and no state estimation is considered to be the most promising and less complex alternative in active control systems in the near future.
1 Introduction Most research on active control systems for structural engineering published since 1980 has dealt with variations on control theory concepts developed in other engineering fields. Linear quadratic optimal control has been the control strategy considered in most published research. The linearity of this controller and phase margin theoretically provided by this technique are the main reasons for its choice. However, the physical realization of such controller in civil engineering structures such as buildings or bridges subjected to earthquake loading implies the use of actuators connected to power sources capable of delivering tremendous power. This requirement typically makes this active control scheme unfeasible because of technological or cost constraints. One exception is the active mass damper system used for wind-induced vibration reduction in tall buildings where the power requirements are much smaller. In fact, more than twenty active mass dampers have been installed in buildings intended for occupancy in Japan during the last two decades. Generally speaking, passive control systems (energy dissipater or reinforcement) that can achieve a desired performance for a given structure will be
301
302 preferred over an active system on the basis of technological simplicity, cost, reliability and maintenance requirements. Semi-active control systems originally proposed in 1974 for suspension of automobiles [1] have been studied since 1980 in the context of civil engineering structures [2,3,4,5]. They constitute a promising alternative for structural vibration reduction because these systems deliver significant power and require negligible external power during operation. A semi-active system consists of an essentially passive device whose mechanical characteristics can be modified in real time using very little power (compared with that delivered by the device) to provide control forces. Recent experimental research on magneto-rheological (MR) fluid actuators in the context of earthquake engineering has shown significant potential of these devices for vibration reduction. Today, MR dampers (MRD) are commercially available under the name of MagneShock™ for automobiles suspension. MRD capable of applying dissipative forces of 20 metric tons have been built and tested for building applications [7]. The dynamic performance of structures with MRD has been investigated experimentally and through numerical simulation [5]. The objective of the research reported herein is to analyze the dynamic performance of structures with variable-damping actuators. Both variable viscous dampers and variable-resistance dampers are considered in the study and compared with the performance achieved through passive damping and fully active systems.
2 Passive and fully active controllers Although several types of energy dissipating mechanisms have been proposed and applied to vibration reduction, viscous fluid and friction dampers are the damping mechanisms used for benchmark analysis herein. The structural models considered are assumed to remain linear during vibration with the effect of supplemental linear or nonlinear components provided by passive dampers and/or actuators. The following differential equations govern the dynamics of the models: My(t) + Cy + Ky{t) + ^LTd / , (t) + £ / £ «,(0 = Lww(t) ,=i
(l)
>=i
M,C,K are the mass, damping and stiffness matrices of the structure, fjft) is a vector of forces in supplemental dampers, u(t) is a vector of forces applied by the actuators, and wft) is the excitation signal. Lj and Lu are transformations relating the degrees of freedom yft) of the structure to the deformations in the dampers Ad(t) and the relative displacement of the actuator Auft), respectively.
303 2.1 Passive damping and LQR controller Generally speaking, better deformation and acceleration reductions can be achieved using passive dampers when the dampers can be positioned within the whole controlled structure. Architectural constraints however, may determine that only a reduced number of devices can be installed and this may impose limits in the reachable performance. It has been proved that linear hysteretic damping is superior than linear viscous damping in the sense of larger deformation reduction for the same maximum force level [4]. It is also known that and increase in energy dissipation in a structure can produce an increase in floor accelerations for large damping ratios, especially in the case of multi-story lightly-damped long-period structures with supplemental dampers installed in a single level. Root mean square (rms) inter-story drift constitutes a simple index for evaluating damage potential a structural system. Let the inter-story drifts be defined as D = LDy(t) (2) Consider a structural model with supplemental linear viscous damping provided by viscous dampers located in a single level of the structural system subjected to broad-band excitation. Let the excitation signal be modeled as a zero-mean whitenoise signal with E[w(t)w(t+T)J=WS(T). The selection of the damping constant c of the dampers can be done by minimizing the trace of the drift covariance matrix with respect to c: PDD=E[DDT] (3) As an example for a benchmark analysis, consider the steel frame structure in Fig. 1 subjected to broad band excitation with supplemental dampers located in the lower level. The lateral displacements yj (?) / = 1,2,...,6 of each floor of the steel-frame are taken as degrees of freedom of the structure. The natural frequencies of the structure are 4.5, 12.5, 21.5, 31.4, 41.1, and 51.5 rad/s. The damping of the structure without dissipaters is modeled as a classical damping matrix C with modal damping ratios of 1% of critical. Figure 2 shows in thick line the square root of the trace of the drift covariance matrix divided by the white-noise excitation intensity, ^] trace (PDD) / W , as a function of c, the damping parameter of the dampers located at the first level of this structure. The circular symbol in the graph shows the optimum damping parameter obtained using this criterion. Although an increase of c above the optimum value produces a reduction in the lower level drift, it produces an increase in the rms inter-story drifts of the upper levels.
304
-+-
-I
[email protected] 1
W24x84
•y6(t) W14x99
W24x; 4
*
W24xl
y,(.t)
* y4(0
W27x: 4
W14xl32
[email protected] W27x 02
W14xl93
Actuator or damper ^
* y2(0
W30) 116
533.4cm
fa
%
^
J3S
Figure 1. Steel frame structure used for benchmark.
Assuming an actuator is installed in the lower level of the structure, full-state feedback is feasible and a powerful actuator is available for design, let us obtain the achievable performance using a linear quadratic optimal control (LQR) with performance index T = E[J x (t)Qx{t) + ru2 (t)dt]
(4)
where £[.]= expectation operator, and the dynamics of the state of the system x(t) is given by (5) x(t) = Ax(t) + Buu{t) + B w(t) withx(0 = LK0
A=
O -M~XK
y(t)Y, I -M'XC
B =
O
B =
O -M'lL.
(6)
and w(t) = white noise signal of intensity W, E[w(t)w(t+T)J=WS(r). Matrix Q in Eq. (4) is selected so as to penalize the sum of the squared inter-story drifts, L'DLD O
fi = o
o
(7)
305
The optimal controller is u(i) = -Gx(t) = -r~'BTuPx{t) where matrix P is the solution of the corresponding Riccati equation.
(8)
Figure 2 shows in thin line, the performance ^trace(PDD)/W achieved as a function of the parameter r of the LQR controller. Figure 3 shows in thick line, the performance for varying c as a function of the normalized rms viscous damper force (fc,) and in thin line, the same performance measure for varying r as a function of the normalized rms actuator force. As Figs. 2 and 3 show, active control can in principle render better performance of the structural system than passive damping in frame structures. However, the marginal improvement is not always that significant and the associated cost may be prohibitive in terms of necessary technology for achieving the force-levels in medium or strong earthquakes.
Iog10(r)
log10(c/[Ns/cm])
Figure 2. Performance of frame subjected to random excitation
log10(RMS(u)/sqrt(W)), log10(RMS(fc)/sqrt(W))
Figure 3. Force requirement of LQR and viscous damping.
As an example, let us compute the response of the steel-frame structure (Fig. 1) subjected to the registered ground acceleration of the El Centra (Imperial Valley, 1940) with no added damping (BF), with optimum viscous damping (VD), with LQR control (r~12, see circular symbol in Fig. 2), and with optimum viscous and LQR control. The results in Table 1 show significant and comparable drift reductions achieved by passive damping, LQR controller and combination of both. It is not surprising that the maximum force in the actuator reaches between 13% and 20% of the building
306 weight, loads that can easily make the physical realization of both active controllers unfeasible for earthquake applications. Table 1. Peak responses of 6-story steel frame subjected to the El Centro signal. Controller
BF VD LQR LQR + VD
Peak force/ trace(Mg) 0 0.148 0.192 0.132
Drift 1 (cm)
Drift 2 (cm)
Drift 3 (cm)
Drift 4 (cm)
Drift 5 (cm)
Drift 6 (cm)
4.47 2.17 2.50 2.39
2.65 1.92 1.43 1.54
3.23 2.62 2.08 2.18
3.62 2.51 2.22 2.19
4.06 2.66 2.81 2.69
3.06. 1.97 2.32 2.11
2.2 Liapunov bounded controller As an alternative to reduce force requirements in an active controller, bounded non linear controllers can be designed using Liapunov techniques. Consider a stable linear system (damped structure) described by Eq. (7). Let the control force be bounded
I "(0 l< «„a,
(9)
Taking the following Liapunov function [4] t u(t)
A(t) = x(tf
Px(t) + 2aj j g~l (-^—)dsdt 0 0
(10)
Mmax
with g(z) monotonically increasing, anti-symmetric and bounded, g(z)=-g(-z) and \g(z)\ < 1 for all z, a Liupunov controller can be designed by minimizing the time derivative of the Liapunov function along the solution of the system with respect to u(t). This yields the following bounded controller u(t) = -umaxg(a-,BTPx(t)) For a stable controller, P is selected as the solution of the Liapunov equation for any positive definite matrix Q. r ATP + PA = -Q
(11)
(12)
To illustrate that this nonlinear controller can achieve deformation reductions comparable to those of an LQR controller but with significantly smaller maximum force requirements, the response of the frame structure with LQR controller and nonlinear Liapunov controller is computed for the El Centro signal. The nonlinear controller is designed as N(t) =
^-tan-" (arxBrPx(f)) n
(13)
307
Q is taken as a block diagonal matrix with LTDLD and the identity matrix / in the diagonal, and umm is taken as a fraction of the building weight. The results are presented in Table 2. NLC 0.05 means umax = 0.05trace(Mg)
where g is the
acceleration of gravity. Table 2. Peak responses of steel-frame subjected to El Centro. Controller
Peak force/ trace (Mg) 0.192 LQR NLC 0.05 0.0470 NLC 0.10 0.0939 NLCE0.10 0.0864 FD 0.03 0.03
Drift 1 (cm)
Drift 2 (cm)
2.50 2.54 2.51 2.54 2.72
1.43 1.65 1.67 1.63 1.99
Drift 3 (cm)
Drift 4 (cm)
Drift 5 (cm)
Drift 6 (cm)
2.08 2.26 2.56 2.28 2.51
2.23 2.62 2.79 2.51 2.64
2.82 3.10 2.97 2.88 2.85
2.32 2.31 2.51 2.36 2.26
The results corresponding to NLCE 0.10 correspond to a particular selection of Q that yields the mechanical energy of the structure. For comparison, the response of the structure with optimum level of friction damping (FD) is also presented in Table 2. For this structure and an excitation model given by the El Centro signal, the optimum yield force in the friction dampers is around 3% of the building weight. As these results indicate, nonlinear controllers with maximum force requirements of 5% of the weight of the structure can render drift reductions comparable to those achieved by LQR with maximum forces of about 20% of the weight. It is worth noting that the nonlinear controllers have not been optimized with respect to a. It is also worth mentioning that friction dampers (FD) can provide performance levels similar to those achieved by nonlinear controllers. 3 Semi-active dampers Two alternative controller designs of semi-active systems are considered in this section: (i) semi-active damper as an actuator for realizing an LQR controller, and (ii) improved hysteresis is SAD through collocated controller design. 3.1 LQR realized through semi-active damper Consider a semi-active viscous damper with controllable viscous constant c(t) (in fact the semi-active viscous damper may show nonlinear dynamic behavior and internal dynamics due to compressibility of the fluid). Consider an LQR controller
308 designed for the structure with optimal control force given by Eq. (10). A feasible implementation of this controller using SAD is c(t) = u(t)/A(t)
if
c(0=0
if
u(t)A(t)0
Table 3 compares the maximum drifts and maximum force requirements of this controller (SA LQR) compared with the corresponding LQR controller for the same simulated earthquake. The results show that clipped optimal control performs as well as the fully active LQR strategy with comparable maximum force levels. While still it may be technologically unfeasible to provide 20% of the building weight with variable dampers, it is by far more feasible than hydraulic actuators realizing an LQR controller. Table 3. Comparison of LQR and clipped LQR controller (SAD LQR) performance
Controller Peakforce/ trace(Mg) 0.192 LQR SALQR 0.233
Drift 1 (cm) 2.51 2.46
Drift 2 (cm) 1.43 1.47
Drift 3 (cm) 2.08 2.10
Drift 4 (cm) 2.23 2.23
Drift 5 (cm) 2.82 2.80
Drift 6 (cm) 2.32 1.91
3.1.1 Can semi-active dampers realize an LQR controller with accuracy ? The ability of SAD to follow an LQR controller can be explained from a probabilistic perspective. A semi-active damper can apply the same control force as the LQR when sgn(GxA) > 0, that is when the LQR control force dissipates energy, u(t) A(t) < 0. It seems appropriate then to analyze the probability of this event for the closed-loop structure controlled with LQR and subjected to a stationary broadband process. Given a structure with LQR controller subjected to stationary Gaussian white noise process, the probability of interest can be computed by transforming to independent unit-variance variables the zero-mean joint Normal random variables u(t) and A(t) with covariance matrix
PIZ=E[[u(t)
A(t)]T[u(t)
A(f)]] =
E[u(t)A(t)] E[u(t)A(t)]
„2
(15)
which can be computed relating z(t)=[u(t) A(t) ]T to the state of the system x(t) and computing the stationary covariance matrix P^ ofx(t). Let's define z(t)=Lzx(t), then P22=LZPXXLTZ
(16)
309 Let O be the orthogonal eigenvector matrix, <J? O = / , and A be the diagonal matrix with the eigenvalues of P K , then A^r r 77 = A 2 0 z = [v(
1 v2Jz
(17)
is a random vector with zero-mean unit-variance Gaussian independent random components. The probability P[u(t)A(t) > 0] can be computed easily in the transformed domain (see Fig. 4) by computing the angle 6 between vectors v, and v,: 0 = cos~'(-
v,||
ii
(18)
and computing the probability as an angle ratio: P[u(t)A(t)0
r) = A2Tz
zi=Gx
Figure 4. Computation of probability in transformed space f]i,T]2.
310 Figure 5 shows in thick line the achieved drift reduction J(r) = -JtraceiP^ ) normalized with respect to the uncontrolled structure J(0), the correlation coefficient and the probability of dissipative control force as a function of the parameter r. For values of r that provide significant marginal drift reductions, the correlation coefficient is larger than 0.95 and the probability of a dissipative control force is larger than 90%. This behavior of an LQR controller will repeat in lightlydamped structures with low and medium level controllers. These high probabilities explain why a clipped LQR controller realized by means of a variable viscous or ER damper can yield performances comparable to LQR controllers in conventional structures. To illustrate a type of structure in which the LQR controller is not dissipative, let us consider two SDOF structures, a main structure and an auxiliary structure coupled by an actuator. The controller is aimed at vibration reduction of the main structural system using the auxiliary structure as a reaction frame. Let the main structure be lightly damped (1% of critical damping), let both structures have the same mass, let the auxiliary structure be damped with 10% critical damping, and let it have a natural frequency three times larger than the main structure. Defining an LQR performance index with a diagonal matrix Q, with coefficients in the diagonal: 1 for the deformation of the main structure and 0.01 for the auxiliary structure, and a range of values of r, the rms deformation of the main structure, J(r)=a , normalized by the rms deformation of the uncontrolled structure, J(0), is computed for the system subjected to white noise support acceleration. Figure 6 shows the results obtained. Not only we see low values of probability of dissipative LQR forces but we also see negative correlation coefficients between A(t) and Gx(t) for strong controllers; this means that a strong LQR controller aimed at energy transfer inputs energy to the system in the mean, something a variable damper cannot do at all. The LQR controller then plays the role of energy transfer rather than energy dissipation in this case. In this type of vibration reduction situation, a clipped LQR controller would perform poorly compared with the corresponding linear LQR controller because with high probability the semi-active damper would not be able to follow the LQR control signal. An even more interesting example of energy transfer can be obtained in a lightlydamped structure with a damped optimum tuned mass damper (TMD) whose performance is to be enhanced by an actuator connecting the TMD to the structure. Again, the LQR controller can yield very low probabilities of dissipative control forces (see Fig. 7) and consequently, can not be realized successfully with a variable damper.
311
• &
1
;orrelo/7
/ /
/ /
Prob[Gxd/dt(A)>0)
J(r)/J(0)
/
Correlation / Coefficient /
/j(*0(0)
y
/
0}
! :
: Correlation Coefficient
— / /
:
Iog10(r)
Figure 7. Probability of dissipative LQR control force in structure with TMD.
These simple results illustrate that in the case of active mass dampers, active controllers can yield significantly better performance than that attainable by passive damping (see J(r)/J(0) in Fig. 7), dynamical systems in which energy transfer prevails over energy dissipation. Although a clipped LQR strategy is not a suitable strategy for vibration reduction in these cases, semi-active dampers can be successfully used in interacting structures using other controllers [8,9].
312
3.2 Variable-friction dampers Several control algorithms have been proposed for variable friction dampers and variable yield-stress fluid dampers (MR dampers). Only work developed by the author on this topic is presented here. Collocated controllers aimed at enhancing the hysteresis loops of the device are presented [9, 10, 11, 12].
3.2.1 Modulated homogeneous friction Modulated homogeneous friction (MHF) [10], a control algorithm for variablefriction dampers or MR actuators, is a collocated dynamic feedback law that uses the deformation of the damper A(f) as the only feedback signal for the control of the yield force of the device: Fy(t) = kd\P[Mt)] (20) where Fy(t) = yield force of the device, kd = positive gain coefficient with units of stiffness and P[A(t)] = prior-local-peak operator, value of the closest prior local maximum or local minimum of the deformation signal. When operated with the proposed control law (Eq. 33), the force in the semi-active damper is given by /[A(0] = kd | P[A(0] | sgn(A(0) (21) The dynamic behavior of structures with MHF actuator has been analyzed in detail elsewhere [10]. Results obtained for benchmarking this controller with LQR are shown in Table 8. The results shown illustrate the high efficiency of the controller in reducing inter-story drifts with relatively low levels of peak force. 3.2.2 Amplitude-modulatedfriction (AMF) In the case of MHF (Eq. 21), the capacity of the friction device was selected to be linear in | H[A(t)] | so as to maintain homogeneity of degree one in the deformationforce relation. Although from an analysis point of view this is a convenient feature that permits very accurate linearization techniques and simplified analysis of the controlled structure [10], from a performance perspective nonlinear relationships between capacity and amplitude of vibration of the type: Fy(t) = h(\P[A(t)\) (22) offer better performance when the yield force capacity of the semi-active damper is limited. In Eq. (22) h(.) is a positive monotonically increasing function for positive
313 argument. For example, if the actuator force is to be bounded by a maximum value F0, the following is a convenient control law: Fy(t)=^tan-l(b\P[A(t)\) y n
< 23 )
With b a positive constant that controls the variation of capacity of the friction device with the amplitude of vibration. Table 4 compares the drift reductions and peak force requirements of MHF and AMF with those achieved using an LQR controller. MHF 0.062 indicates the assumed value of kd I K{\,\) = 0.062, AMF 0.05 indicates that F0 ltrace{Mg) = 0.05. As in the case of passive friction, an increase in the yield capacity may turn out to be deleterious, producing an increase in the inter-story drifts of the higher levels. Since the device is operated as an enhanced friction damper, the distribution of the variable-resistance dampers over the height of the building is recommended. Table 4. Performance of structure with MHF. Controller LQR MHF 0.062 AMF 0.05 AMF 0.10
Peakforce/ Trace(Mg) 0.192 0.034 0.049 0.097
Drift 1 (cm) 2.51 2.51 2.60 2.39
Drift 2 (cm) 1.43 1.79 1.94 2.09
Drift 3 (cm) 2.08 2.37 2.58 2.62
Drift 4 (cm) 2.23 2.50 2.54 2.38
Drift 5 (cm) 2.82 2.98 2.83 2.83
Drift 6 (cm) 2.32 2.28 2.00 2.32
3.2.3 A new modulated homogeneous friction (MMF) controller The following aspects motivated the development of the MHF controller presented above: (i) simplicity of the controller, (ii) homogeneity of degree one so as to ease the design, and (iii) rectangular hysteresis loops to maintain the efficiency of passive friction dampers that maximize the energy dissipation per cycle for a given maximum force applied to the structure. Another motivation in the creation of this controller was to keep in the semi-active damper, a rate-independent behavior following the fact that linear hysteretic damping offers better performance than linear viscous damping in terms of deformation reduction and maximum accelerations in the controlled structure [4]. Although a rectangular hysteresis loop shows maximum efficiency in terms of energy dissipation per cycle for a given maximum force in the damper, it is not
314 necessarily the most efficient shape in terms of maximum base shear. The reason is that it is the sum of the elastic force and the dissipative forces what produces the maximum shear on the foundation of the structure. It is apparent that for a SDOF structure with dampers that provide rectangular hysteresis (such as dry friction), the maximum shear will coincide in time with the maximum contribution of elastic forces to the base shear which in turn coincides with peaks of the deformation signal. Thus, a new MHF controller is proposed that reduces the damper force as a function of the current deformation, so as to reduce the maximum total force transmitted to the adjacent floors of the structure in the vicinity of local peaks of the deformation signal. The proposed new MHF algorithm is
Fy (t) = kd | P[A(t) 11 (1 - y I - j ^ - 1 ) I // P[A(0] * 0, | P[A(0 |>| A(/) | L WJ Fy(t) = 0 ifP{A(t)]=0, or | P[A(t) \
Specimen
f
Computer A for sating
r~i
• StnmCtga ^
Oscilloscope
J
Strain Meter
Computer A for rasing
|
• SmuiOage
Osciucscope
*—
f
L.
r
DC Amplifier
i
/ l
L
Input (Shaking Title)
_.. ,, uuiu,
Computer B far input
e a : Strain Gage
\— i—1
• •
: Piezoceramics
- Piezoceramics V )
i — >
J 1I
"N
OsdllUar for input
Transfonncr
input (Shaldnn Table)
t = j : Strain Gage
mt
> Computer B for input : Piezocenmics
F i g u r e s 6 (a) & (b): Systems for evaluating sensing function and evaluating actuating function
r^
Mass(Lead
-
JL ^Fb
Spodmoa
-65X«
ComputerA ftrKoiing
S t r a i n Gage iaae liable Rautor
. - S t r a i n Gage /
21 -150x150x1
il
StaunMeter • Oscilloscope
• 2L -50x50X4
BlmorpJ
in /Variable Reaisrot
CD
3 Shaking Table
7777 ^ ^ ^ i ^ k t a g T a b l e ///>///
•SmJiiOqe
B a : StninGage
Computer B for input
• * : Piezoceramics
F i g u r e 7: Portal frame specimen
F i g u r e 8: System for evaluating damping
(Type B specimen)
function
In all of the tests the excitation was done with a shaking table. For the sensing verification tests the excitation forms were sinusoidal excitation with, f, of 6.0 Hz, and ll.OHz, and for the actuating and damping verification tests, excitation was four sinusoidal cycles at the system's natural frequency, / , of 6.0 Hz and 12.8Hz. In actuating verification testes, an AC voltage of 250V, and an inverse phase of 6.09Hz was applied to the piezoelectric materials. The damping ratio, h, in each test was calculated from vibration data measured for free vibrations, by using a least-square method based on the following equation:
345 |ea| = anexp{-/i~-r}
W
Where | e l is the strain amplitude, / is time, a is the amplitude at ?=0, and T, is the natural period. 4.2 Optimum resistor in piezoceramic damper Under the steady-state vibrations, the equivalent damping ratio added by the piezoceramic damper, ft , is obtained as follows[5,6] fc-=2A 2 v Where,
"
(l-*?l)+p p = R-c{l-kll)-2n-f
(2)
(3b)
77: Loss factor, p : Non-dimensional frequency, V : Peak strain energy in the piezoelectric materials and the total system. fc3]: Electromechanical coupling coefficient ( 3: Polling direction: 1: Vibration dirction), R : Resistance {€!), C : Static capacitance (F), / : Frequency of excitation The optimum resistance, R t e that maximizes ft can be determined as ^s«. = o 3R Consequently,
R D P ,(/)=- 7 f =4i—
(4)
(5)
346 4.3 Results and discussion (a):(f=6.0Hz)
(b):(f=11.0Hz)
Figure 9: Strain indicated by piezoceramic sensor (a)with piezoceramic actuator
(b)without piezoceramic actuator
Figure 10: Time history of strain response of type A specimen compared with strain indicated by strain gage (a) 0.0 to 10
(b) 0.0 to 2.0
1.2 1
ku-tJi UJH.U7
yMwdmm
1
$0.1 . w
Ak
0
f X. *" 0
-.^
4 0,4
a
,
•
km-tJl ViflXW
Ucdmm
Exptrimm .
: - w - « -JL "*"
000
"***"•"—~~i
°v.o,o:
—'W*t»
< fi.5 /
OpttMimRistHiKt
-.y..R^mi$aj^
'
"X![ty-titoO ) 0
0.2 i
Exfrimta 4-(
0 5
10
0
0.5
1
1.5
Figures 11 (a), (b): Measured and calculated damping ratios for R/R varied from (a) and (b)
347 Figures 9 (a) and (b) show results for sensing tests, Figs. 10 (a) and (b) show results for actuating tests, and Figs. 11 (a) and (b) show results for damping tests. The temporal strain response of the piezoelectric materials are shown in Fig.9 (a) for/=6.0Hz and in Fig.9 (b) for/=11.0Hz. Figures 10 (a) and (b) show the temporal vibration of strain in the vicinity of the fixed end with and without actuation, respectively. In Fig. 10, £ is normalized by the strain, £a measured when the top of displacement of the specimen reaches. l.O(cm) and t is normalized by the natural period, ;/T, . Figure 11 shows R/R m , vs h, where Row is calculated from Eq.(5) and h is calculated from Eq. (1) The result shown in Fig.ll (a) correspond to R/R0IU was varied from 0 to 10 and the results shown in Fig.ll (b) correspond to R/Rou, varied from 0 to 2. In Figs. 11 (a) and (b), the relationships between R/Rou; and h^ + hinh calculated from Eq.(2) are also plotted. Where, the inherit damping ratio, hinh , is identified as the experimental observation when the resistance is to be 110(k£2). From these results, we clarified that 1) The amplitude and phase of strain measured with the piezoceramic sensor depended on the frequency of excitation. 2) The piezoceramic actuator was able to control the vibration at 1.5 times its inherit damping ratio. 3) The optimum resistance of the piezoceramic damper that maximized the damping ratio was accurately predicted by Eq (5). 4) When the damping was attached over only 10% of the column, the piezoceramic damper increased the damping ratio by 30% compared with its inherit damping ratio.
5.
Prospects for Smart Material Systems as Future Trends
A smart material is one that not only adds intelligent function, but one that also functions as a sensor, data processor, actuator, and expression for external disturbances. That is, smart materials add an aspect of artificial life. This is different from intelligent materials, which only respond according to a single set of inputoutput (non-evolving) response characteristics. Recently many research studies are being made on artificial life, focusing on evolution, shape formation, learning, distributed parallel biological processing, immunity, and self-remodeling [7,8]. As one of the basic mathematical functions of artificial life, genetic algorithms based on the principle of biological evolution (i.e., selection, crossover, and mutation) are the models for the evolution process. Among the possible processing functions, evolution is the most useful method for optimization, because the system responds
348 according to simple internal principles and through interactions with outside sensors, and not by external instructions. Thus, self-organization can independently form the system order. Also, self-formation is the mapping from a genetic type to an expressing type, and has the important role of enhancing the robustness of the system adaptability. In designing and controlling large-sized, complicated response-control systems for buildings that are in uncertain and changing environments, it is impossible to provide control algorithms and data that can handle every control scenario. As the complexity of the control system increases, the possibility of providing accurate, comprehensive information decreases, thus degrading the responsiveness, reliability, safety, and robustness of the system. To avoid this, the development of smart material systems that use them is needed. Figure 12 shows the overview of smart material systems that use them, as modification to reference [9]. External Diitwbance | .,— •„,„ jv f
Controller
r Evolution, Learning, Monitoring and Repairing, Shape Fomiatian, Immunity, Self-Remodeling, and Distrib. uted Parallel Processing Figure 12: Overview of smart material systems (added to 9) Figure 13 shows the technical development of building structures from the late 20th century to the first half of 21 st century. Currently, earthquake disaster countermeasures for buildings are seismic design and response-control devices, such as actuators and sensors. Monitoring is also a must for maintaining and controlling such response-control devices. In the future, incorporating genetic algorithms into response-control systems will make the self-organization of systems and their organic optimization economically possible (i.e., economization). Fuzzy theory and neural networks are examples of artificial intelligence. It is possible to merge the characteristics of each of these algorithms by adopting genetic algorithms for selecting or improving the rules of fuzzy theory or neural networks. Consequently, introducing smart material systems into buildings is practicable only
349 when coupled with the development of smart materials. Structural interpretation of the aspects of artificial life (i.e., evaluation, shape formation, learning, distributed parallel processing, immunity, and self-remodeling) will coincide with the development of technology related to each aspect. Incorporating smart material systems into buildings (i.e., smartization) is the future of earthquake countermeasures; the ultimate goal of engineers is to design a buildings that behave like a human being.
( Lite20lhCentury) (At Present) Seismic Design Reipoiue Control
! First Half of the 21sl Century) (IniheFmurc)
~~* Mechanization (Inorganic)
[
Development of Smart Material
)
MoolMrlDg
\
( Setf-Onnniatloii ^ v ^
(Eamorabatlon j ArttflclaiLlfe (Organic) Environmentally-Destructive Structures Eletooilary Trentoent of System Evaluation from the Viewpoint of Construction planning (Strength Capacity, Durability, etc.)
>* r
Dtvtkpmeat of Smart Structures (l.t, Srantitaton of Structures)
T Environmentally-Friendly Structures Treatment of Total system Evaluation from the Viewpoint of Maintenance (Life Line, Exchange Parti, etc.)
F i g u r e 1 3 :Post and future technical development o f building construction
6.
Concluding Remarks
In this paper, we first summarize the control principles behind structural responsecontrol systems, and also reviewed the required performance of control systems. To demonstrate how piezoelectric materials can be incorporated into smart material systems, we made vibration tests with scaled cantilevered beam and portal frame. The results of vibration tests confirmed the validity of the control effect of the response-control systems and demonstrated the possibility to clarify the characteristics and performance of piezoelectric materials for use as actuators, sensors, and dampers in a smart material systems. Finally we speculated on the general concept and the prospects of using smart material systems in future building designs. The performance of structural response-control systems depends on the control devices, which have been developed mainly from a practical-use viewpoint.
350 When we reach a consensus on the requirements of future structural control systems, incorporation of smart materials into these systems will proceed. When that occurs, it will be important to develop reliable smart material technology and methods for evaluating the vulnerability of each component of structural responsecontrol systems. References
1.
2.
3. 4. 5.
6. 7. 8. 9.
BRI Technical Report,"R&D to achieve active response-control structurescooperative study between public organization and private company",MOC (1992, 1993, and 1994). Fujitani, H., Midorikawa, M., Iiba, M., Kitagawa, Y. et a\."Seismic response control tests and simulated by fuzzy optimal logic building structure", Engineering structures, vol.20, No.3. (1998). Report of project on R/D to achieve structural response-control of building structures, MOC, Government of Japan, (1993, in Japanese). Kitagawa, Y., Tamai, H., and Takeshita, M. "Smart structural systems of exposed to external, disturbances-concept and technology'', 12 WCEE (2000), CD-ROM Hagood, N.W., Flotow, A. Von "Damping of structural vibrations with piezoelectric materials and passive electrical networks", J. Sound and vibration, vol.146, No.2,(1991),pp.243-268. Hagood, N.W., Crawley, E.F "Approximate frequency domain analysis for linear damped space structures", AIAA Journal, vol. 28. (1990), pp.1963-1961. Labgton, C. Artificial Life, (1989), Santa Fe. Kitano, K.,"Artificial life and combination of evolution, generation and learning", Mathematics Science vol. 353. (1992) Report on "Leading Research R&D of Smart Structure Systems", NEDO-PR-95011, (1996).
THE MOST RECENT APPLICATIONS OF SEISMIC ISOLATION AND PASSIVE ENERGY DISSIPATION
ALESSANDRO MARTELLI Chairman, Working Group on Seismic Isolation (GLIS) of the Italian National Association for Earthquake Engineering; ENEA, Bologna, Italy E-mail martelli @ bologna, enea. it MASSIMO FORNI Technical Secretary, GLIS; ENEA, Bologna, Italy E-mail forni@bologna. enea. it Summarized in this paper is the state-of-the-art on seismic protection through innovative antiseismic techniques, namely seismic isolation (SI), passive energy dissipation (ED) and active control (AC), based on the information collected at the Post-SMiRT Conference Seminar on Seismic Isolation, Passive Energy Dissipation and Active Control of Vibrations of Structures held at Cheju (Korea) in August 1999 and on even more recent information which became available to the authors. Reported is information on the most recent applications of such techniques, together with the progress of R&D activities at word-wide level, availability of design rules and the related issues and needs for further activity. With regard to the latter, somewhat more detailed information is reported for European and especially, Italian applications.
1. Introduction Modern society is being more and more characterized by a strong interaction among the large systems by which it is formed: the physical, human and infra-structural systems. Seismic risk results from the interaction among seismic hazard, vulnerability of structures and social-economical effects. In the past, an earthquake mainly caused collapse of buildings and fatalities. Nowadays, a seismic event may also endanger the social-economical stability of large areas, due to the complexity of technologically advanced societies. For instance, the Great Hanshin-Awaji earthquake of 1995, which struck Kobe (where one of the most important ports of the world is located) is the first case in the history of a seismic event that occurred in a highly industrialized urban area, by producing enormous damage to the building, road and in particular, productive systems. The earthquake which struck Izmit in Turkey on August 17, 1999, caused the fire of the biggest Turkish petrochemical plant, by leading to very difficult fuel supply and heavy pollution. A scenery similar to those mentioned above might take place in many other areas in the world, from California to Italy: in California, for instance, in case of a strong earthquake closer to San Francisco and Silicon Valley, 351
352
with the respect to the 1989 Loma Prieta event; in Italy, for instance, in case of events like that which struck the now highly industrialized area around Po River in 1117, or that which destroyed South-East Sicily (where a huge number of petrochemical plants and components is now located) in 1693. In addition, it is worthwhile mentioning that recent earthquakes showed a fully unexpected violence, like for instance, that which struck again Turkey, with epicenter near Kaynasly (Bolu Mountains), on November 12, 1999. Since ground acceleration was much larger than the design value, this caused severe damage even to some very important modern structures, like a viaduct of the new IstanbulAnkara freeway, being erected using the most modern anti-seismic technologies, which was extremely close to the epicenter [1]: in fact, the maximum displacement allowed by the horizontal fail-safe system (stoppers) was largely exceeded (another viaduct behaved very well, although it displaced twice the design value, but this was still allowed by the stoppers). The aforesaid remarks demonstrate, without any doubt, the increased degree of complexity of modern society, and thus, the need for an integrated management of the territory, able to make development and safety compatible. This implies that more and more numerous shall be the structures for which design shall not be limited to prevent their collapse, but shall require the absolute integrity and full operability after the earthquake. The feature of absolute integrity is also indispensable to protect investment, taking into account that the value of contents of more and more buildings is much larger than that of the structural members, as well as to avoid spending the enormous amounts of money during both the emergency phase and reconstruction which were necessary after the recent earthquakes. For the above-mentioned reasons, a wide extension of the use of innovative anti-seismic techniques, such as seismic isolation (SI) and passive energy dissipation (ED), which aim at ensuring the full integrity and operability of structures, is necessary for both new constructions and retrofit of existing buildings [2]. In fact, SI and ED technologies are now fully mature for such an use, as demonstrated by the results of very numerous research projects and also, by the excellent behavior of seismically isolated buildings in both the Great HanshinAwaji earthquake and the Northridge earthquake which struck the Los Angeles area the year before [2]. This conclusion has been confirmed in all the recent Conferences on seismic engineering, in particular at the 1999 International Post-SMiRT Conference Seminar on Seismic Isolation, Passive Energy Dissipation and Active Control of Vibrations of Structures [3]. 2.
Recent applications
The invited lectures and contributed papers presented at the Cheju Seminar and the
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extensive discussion both following their presentation and during the Closing Panel, demonstrated that not only SI, but also several ED systems are already fully mature for wide-ranging applications. They also showed that, at last, the benefits of such systems have been well understood in several countries and that they are now being more and more used. The aforesaid benefits had already been very well understood by Japanese after the 1995 Kobe earthquake and to a certain extent, by Californians after those of Loma Prieta (1989) and Northridge (1994). Even before, this had occurred in New Zealand, where there are still new applications of SI to both new and existing ancient constructions, in spite of the limited population; more recently, it also occurred in other countries, like the P.R. China, Russian Federation (especially after the Sakhalin earthquake in 1994) and Italy (after the earthquake that struck Umbria and Marche Regions in 1997, by severely damaging famous frescos of Cimabue and Giotto in the "San Francesco Basilica Superiore" at Assisi). It is also worthwhile citing again that, according to the information provided at the Cheju Seminar, SI and ED are now considered of great interest also for areas characterized by low or moderate seismicity [3]. 2.1 Applications in Japan In Japan the number of buildings provided with innovative anti-seismic systems is still considerably increasing, in spite of the need for still asking for a specific approval for each design including these techniques [3]. The number of licenses began to drastically increase in September 1995, some months after Kobe earthquake (60 new applications) and the annual number reached 207 in 1996, while the overall number during the 10 previous years was 79; such a dramatic increase ended in 1997, when a probably steady progress began (the new licenses were 135 in 1997 and 131 in 1998). In this country, the use of SI was recently extended from new constructions to retrofit of existing buildings (e.g. Le Courboisier Museum at Tokyo), as well as to many new or existing bridges and viaducts (in some cases, at least in Kobe, becoming compulsory for the latter). SI is having many variations in application objects, application methods of rubber bearings and kinds of devices. The variations include SI of tall buildings of about 100 m height, SI of artificial grounds for multiple buildings, and application of non-rubber type SI devices. SI is finding new applications which include wooden houses, masterpieces in museums, automatic storage systems of warehouses, etc.: for wooden houses, nonrubber type SI systems using ball/rubber bearings or sliding bearings to support the superstructure have been developed and began to be used; for masterpieces in museums, various types of SI systems have been developed and used for the containing showcases; for automatic storage systems of warehouses, a new type of SI floor has been developed and used.
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It is also noted that SI is beginning to be used for very important public buildings and facilities, such as, for instance, the new official residence of the Japanese Prime Minister. 2.2 Applications in the USA In the USA (especially in California), new constructions of important isolated strategic buildings, including emergency control centers (e.g. those at San Francisco and Long Beach) are going on and retrofit of even large public buildings using SI (e.g. the San Francisco City Hall and San Bernardino Medical Center) is progressing [3]. Most applications make use of rubber bearings, namely High Damping Rubber Bearings (HDRBs) or Lead Rubber Bearings ( LRBs). However, the extent of the aforesaid progress is much less than in Japan. In fact, although the first US seismically isolated building was completed in 1985, in 1999 there were in this country only 25 applications to new constructions and 22 retrofits of existing buildings: this is due to very complex and conservative regulations. Conversely, SI is now being widely used in the USA for highway bridges, for which it is governed by a simple and not overly conservative code. 2.3 Applications in New Zealand There were 10 isolated buildings in New Zealand in 1999 (in addition to several applications to bridges and viaducts), four of which being retrofits of ancient constructions (those of the Old Bank of New Zealand and Wellington Museums were completed in 1999) [3]. Most applications make use of LRBs, in some cases in conjunction with teflon sliders. 2.4 Applications in Other Non-European Countries As regards other non-European countries, in the P.R. China there were already 160 buildings isolated by means of rubber bearings in 1999 [3]; the total numbers of Chinese isolated buildings and bridges & viaducts reached 230 and 20, respectively, in May 2000 [4]. In Taiwan 10 bridges had been supported by LRBs, in addition to others being erected using viscoelastic devices (VEDs) and elastic-plastic (EP) dampers [3]. It is also worthwhile noting that the number of seismically isolated bridges using LRBs was approaching 30 in Korea in 1999, in spite of its low and moderate seismicity [3]: the main reason is that the use of SI is generally accepted in Korea as an alternative way to reduce the additional construction costs caused by the seismic design requirements recently adopted in this country. With regard to important new applications of ED systems to bridges and viaducts, to be cited are also those to: - three new viaducts of the Istanbul-Ankara freeway in the Bolu Mountains (Turkey), two of which completed and one under construction, which have been
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provided with 'multidirectional EP devices (Figure 1): as previously mentioned and explained in [1], two of these viaducts behaved in an excellent way in the earthquake of November 12,1999; - the Bangabundhu Bridge over the Jamuna River in Bangladesh, with hysteretic devices [5, 6]; - 26 important railway viaducts in Venezuela, again with hysteretic devices [5, 6]; - 5 bridges along the North-South Route in Chile, with VEDs [5, 6]. In Chile, the new hospital of the Catholic University has been isolated with HDRBs [5, 6]. 2 J Applications in Western Europe As regards Europe, to the knowledge of the authors, most new applications of the innovative anti-seismic techniques (in progress or under design or planned) concern Italy, where there were already over 30 applications of such techniques in 1998 [2], 2.5.1 Applications in Italy. In Italy to be cited are the following recent / new applications to [3, 5, 6]: - The "San Francesco Basilica Superiore" at Assisi (Umbria), which had been severely damaged by 1997 earthquake: in October 1999, it was equipped with Shape Memory Alloy (SMA) devices and innovative shock transmitters (the latter developed in the EC-funded REEDS Project, in the framework of the restoration of the Basilica (see Figure 2 and [7]). - The "San Giorgio in Trignano" Bell Tower at San Martino in Rio (Reggio Emilia, Emilia-Romagna), which had been severely damaged by the Reggio Emilia and Modena earthquake of 1996: in November 1999, it was also retrofitted using SMA devices, in the framework of the EC-funded ISTECH Project (see'Figure 3). - The San Feliciano Cathedral at Foligno, damaged by the 1997 quake, being retrofitted using SMA devices. /--CONTINUOUS SLAB
Figure la. Viaduct N. i of the Istanbul-Ankara Figure l b . Detail of the pier top of the Viaduct of freeway provided with EP devices. Figure la.
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Figure 2 a. SMA devices installed on the ''San Figure 2 b. Shock transmitters installed between Francesco Basilica Superiore" at Assisi (PG). nave and transept of the "Basilica Superiore".
- The "La Vista" and "Domiziano Viola" schools at Potenza (Basilicata), which were retrofitted in 1999 using dissipative braces (see Figure 4). - The "Gentile Fermi" school at Fabriano (Marche), a reinforced concrete building constructed in the years s50s, being one of the few examples of rationalist architecture in the town, which had been also heavily damaged by the 1997 earthquake: it is being retrofitted using VEDs (see Figure 5). - An apartment building, under construction with HDRBs at Rapolla (Potenza, Basilicata) close to a twin conventionally founded building (see Figure 6); this application is similar to the twin isolated and non-isolated buildings already existing at Squillace (Catanzaro, Calabria) [2]. - The "Rione Traiano" Civic Center at Soccavo (Naples, Campania), a very large consfruction erected with conventional foundations before the 1980 CampanoLucano (Irpinia) earthquake, when the area was not considered as seismic: this is being retrofitted using approximately 500 HDRBs. - A new hospital at Frosinone (Lazio), which has been designed using HDRBs. - Several buildings of the new Emergency Management Center for Central Italy at Foligno (Perugia, Umbria), being designed using various innovative anti-seismic systems (see Figure 7). - The new hospital at Perugia and some apartment buildings at Citti di Castello (Umbria), being isolated with HDRBs.
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Figure 3. Application of SMA devices on the Sail Giorgio in Trignano (RE) bell-tower.
- Two new buildings at the Navy Base of Augusta (Siracusa, Sicily), to be probably isolated using HDRBs, similar to those already existing at such a Base [2]. - An electric substation at Laino (Calabria), to be isolated by the Italian Electricity Board (ENEL) with wire ropes, based on the results of an extensive numerical and experimental study (this will be the first electric equipment in Italy to be provided with a SI system). - Several viaducts of the Salerno-Reggio Calabria freeway (Campania, Basilicata and Calabria), for which retrofits using ED systems are being designed. - A bronze statue of Germanicus Emperor, located in a museum at Perugia (Umbria), which was provided with a multistage SI system using HDRBs (this will be the second Italian application of SI of this kind, following that to the famous Bronzes of Riace at the Reggio Calabria Museum [2]).
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Figure 4. Dissipalive braces installed on the "La Vista" and "Domiziano" schools (Potenza).
Figure 5 a. The 'Gentile Fermi' school at Fabriano damaged by the 1997 Umbria&Marehe earthquake. Cut of the walls for the introduction of the breaces supporting the viscoelastic devices (Figure 5 b).
Figure 5 b. Viescoelastic device (REEDS Project) similar to those installed on the Fabriano school.
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Figure 6. Isolated under construction and already erected conventional buildings at Rapolla (PZ); [IDRB during the installation phase.
Figure 7. Sketch of one of the buildings of the Center of the new Emergency Management Center of Central Italy, Foligno, Umbria, which will be seisrnically isolated using HDRBs.
In addition, SI might be adopted in Italy for other buildings or structures; in particular, based on the already promising results of an ongoing study funded by the National Group for the Defense from Chemical, Industrial and Ecological Risks of the National Research Council (CNR), it may be adopted for Liquefied Natural Gas (LNG) tanks, such as an existing spherical butane storage tank located in a highly seismic Italian site, as a possible pilot application in Italy for chemical plants (Figure 8) [8].
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Figure S. Sketch and Finite Element Model of the IMG tank selected as possible pilot application.
Finally, the possibility of reconstructing ancient villages in Marche and Umbria Regions using the original masonry materials and to make it feasible, SI is being considered in Italy: to this aim, under consideration are the village of Mevale di Visso in Marche Region (which was almost fully destroyed by the Marche and Umbria earthquakes of 1997-98, after being severely damaged by previous earthquakes) and villages around Nocera Umbra in Umbria Region (which were also severely damaged by the 1997-98 earthquakes). Data concerning the Italian applications are available on Internet at the GLIS address: http://192,107,65.2/glis. It is noted that some applications, in particular those of the new Emergency Management Center at Foligno and (if confirmed) that to the reconstruction of Mevale di Visso, will take advantage of collaborations recently established (or being established) between ENEA and Italian Regions for carrying out pilot applications on buildings, by joining the use of innovative anti-seismic systems with the energetic-environmental quality [5, 6, 9]. 2.5.2 Applications in Other Western European Countries. With regard to other Western European countries, new / recent important applications known to the authors are to [3]: - Two storage tanks of Lonza Company for hazardous chemical materials, which were retrofitted at Visp (Switzerland) using HDRBs. - Bridges in France, including a "TGV" fast train bridge at Marseille, which were provided with VEDs. - A French building at La Martinica to be provided with VEDs in conjunction with cables. - The 4 km long Santarem cable-stayed bridge over Tagus river, isolated using
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HDRBs (Portugal). - The "21th April" suspension bridge over the Tagus river (Portugal), which was upgraded using viscous dampers (VDs). - The new "Vasco de Gama" Tagus crossing (Portugal), which was provided with shock transmitters, VDs and elastic-plastic (EP) devices. - Some small bridges in Greece, provided with HDRBs. 2.6 Applications in the Former USSR Countries Some new building applications of SI were also carried out in the former USSR countries (where the total number reached 306 in 1999 [3]: these were performed in Russia, Armenia and Uzbekistan and made use of HDRBs (the previous applications mostly made use of so-called "low cost isolators"). To be cited among the aforesaid recent applications of HDRBs is the retrofit of the bank of Irkutsk-City (Russia), where isolators manufactured in the P.R. China were installed. 3.
State-of-the-art on R&D
The papers presented at the Cheju Seminar also showed that most necessary R&D activity has already been completed, not only for SI, but also for most types of ED systems (further work remaining necessary for very new devices such as electromagnetic dissipators [3]. With regard to studies performed in the European Union on SI and ED systems, those previously mentioned, concerning the REEDS and ISTECH Projects for the optimization of hysteretic, viscous and viscoelastic dampers and shock transmitters, as well as the development of innovative rolling SI systems and SMA devices, had been just completed at the time of the Seminar and confirmed the excellent behavior of such devices [3]. To be cited is also the present availability, of test equipment - not only in Japan, but also in the USA (for instance, that of Caltrans at San Diego) - capable of qualifying full or at least, large scale devices, as necessary to correctly estimating safety margins: in fact, such tests, if performed on small scale devices, may be not very satisfactory even for rubber bearings (for instance because bonding conditions could be different from those of the real scale device) and are certainly not adequate for VDs, because they cannot correctly describe fluid heating conditions [3]. Regarding future studies and application of SI and ED systems, issues which were stressed were the importance of: - extending retrofit using the innovative anti-seismic techniques; - improving studies concerning innovative systems applicable to cultural heritage; - improving knowledge and develop systems for vertical isolation; - promoting more applications to hospitals and chemical plants and components;
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- widely extending application from strategic to apartment buildings; - performing adequate monitoring; - improving knowledge on seismic input, in particular for near-field earthquakes (how correct is this point was confirmed later by the aforesaid earthquakes in Turkey); - improving studies concerning some reliability and uncertainty issues which have not been yet fully analyzed (including scale effects for qualification tests of SI and ED devices, the behavior of such devices at earthquake levels exceeding the design value, and failure modes, at extremely violent beyond design earthquakes, of structures provided with the anti-seismic systems); - considering other sources of vibrations which may damage or weaken structures, for instance, traffic. Finally, with regard to non-passive control systems (active, semi-active and hybrid systems), the papers presented by the experts at this topic stressed that also their development is further progressing well. Thus, it was decided that the attention devoted at Cheju to this topic has to be kept also at the next Seminar. 4.
Design guidelines development
The only still remaining problems for a wide-ranging application of SI and ED systems that were stressed at the Cheju Seminar concern the design rules for structures provided with such systems [3]. In general, the situation did not improve much with respect to the previous Post-SMiRT Conference Seminar held at Taormina in 1997 [10], especially because such rules are still different in the different countries, frequently still penalize the use of SI with respect to the conventional design and their application still requires heavy approval processes. The only important improvement is that there are now, at least, design guidelines available in most countries (including Italy, where they were only very recently published by the Ministry of Construction). An interesting recommendation made in the Closing Panel of the Cheju Seminar was to try to find the way to develop international design guidelines for structures provided with the innovative anti-seismic systems. Among others, these international guidelines should explain such systems correctly and leave official codes out of consideration. They would not have any legal value, but may be useful, because they would be based on knowledge and experience of real experts. This guidelines' development might be part of the activities of the International Earthquake Research Center that had been proposed at the main SMiRT Conference, held at Seoul (Korea) the week before that of Cheju Seminar. The problem to allow for these activities is to find the necessary funding. In the aforesaid Closing Panel, it was also proposed the guidelines of all countries represented in the Seminar to be translated in English and published in an
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appropriate volume at the next venue, and that, at such a venue, there shall be papers on applications, each containing sufficiently detailed reference to the codes used in the related country. With regard to non-passive control systems, it was stressed at Cheju that the development of these techniques suffer from the fact that they are not considered at all by design rules [3]. 5.
Conclusions
Based on information collected at the International Post-SMiRT Conference Seminar on Seismic Isolation, Passive Energy Dissipation and Active Control of Vibrations of Structures, held at Cheju (Korea) in 1999 and more recent information that became available later to the authors, the state-of-the-art on the applications of SI and ED systems has been shortly reported and some remarks on the progress of R&D activities at word-wide level and design guidelines development have been made. It has been stressed that SI and ED technologies, which aim at ensuring the full integrity and operability of structures, are fully mature, as demonstrated by both the results of very numerous research projects and the excellent behavior of seismically isolated buildings and viaducts in violent earthquakes. It has been shown that, consequently, a wide extension of the use of these techniques is in progress, for both new constructions and retrofit of existing buildings. References 1. Marioni, A., The Effects of Recent Earthquakes on the Base Isolated Bridges of the Istanbul Ankara Motorway near Bolu. Proceedings, IASS Symposium on Bridging Large Spans - From Antiquity to Present, Istanbul, Turkey, (2000) pp. 261-270. 2. Martelli, A., and Forni, M., Seismic Isolation of Civil Buildings in Europe. Progress in Structural Engineering and Materials, Construction Research Communications Ltd., London, 1 (3), (1998) pp. 286-294. 3. Koh, H.M., and Martelli, A., Preface - Overview and Summary of the International Post-SMiRT Conference Seminar on Seismic Isolation, Passive Energy Dissipation and Active Control of Vibrations of Structures. Seismic Isolation, Passive Energy Dissipation and Active Control of Vibrations of Structures - Proceedings of the Post-SMiRT Conference Seminar, Cheju, Korea, August 23-25, 1999, Seoul, Korea, (2000). 4. Zhou, F.L., Recent Development on Isolation and Energy Dissipation Used in New Seismic Design or Retrofit for Structures in China. Proceedings, IASS Symposium on Bridging Large Spans - From Antiquity to Present, Istanbul,
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Turkey, (2000) pp. 230-240. 5. Martelli, A., and Forni, M., State-of-the-Art on Seismic Protection through Innovative Techniques. Proceedings, IASS Symposium on Bridging Large Spans - From Antiquity to Present, Istanbul, Turkey, (2000) pp. 251-260. 6. Martelli, A., and Forni, M., State-of-the-Art on Recent Applications and Research Needs, Session on Base Isolation System. Proceedings, Final Workshop on "Protezione Sismica dell'Edilizia Esitente e di Nuova Edificazione attraverso Sistemi Innovativi; Programma MURST PRIN 97, Naples, Italy (2000). 7. Castellano M.G., and Martelli, A.,The Influence of Shape Memory Alloy Ties on the Seismic Behaviour of Historical Masonry Buildings. Proceedings, IASS Symposium on Bridging Large Spans - From Antiquity to Present, Istanbul, Turkey, (2000) pp. 271-280. 8. Forni, M., Martelli, A., Poggianti, A., Spadoni, B., Pugliese, A., Sano, T, Ciampi, V. and Foraboschi, F.P., Development of Innovative Anti-seismic Passive Systems for the Protection of Industrial Structures and Components. Proceedings, Second European Conference on Structural Control, Champs-surMarne, France (2000). 9. Martelli, A., Forni, M., Bettinali, F., Bonacina, G., Bergamo, G., Castellano, M.G., Medeot, R., Marioni, A., Sano, T., and Pugliese, A., New Activities Performed in Italy on Innovative Anti-Seismic Techniques for Civil and Industrial Structures. Proceedings, 1999 ASME-Pressure Vessel and Piping Conference, Boston, Massachusetts, USA, PVP-Vol. 387, ASME, New York, (1999) pp. 311-326. 10. GLIS, Seismic Isolation, Passive Energy Dissipation and Active Control of Seismic Vibrations of Structures - Proceedings of the International Post-SMiRT Conference Seminar, Taormina, Italy, August 25 to 27, 1997. A. Martelli and M. Forni eds., Bologna, Italy (1998).
THE STATE OF THE ART IN STRUCTURAL CONTROL IN ARMENIA AND PROPOSAL ON APPLICATION OF THE DYNAMIC DAMPERS FOR SEISMICALLY ISOLATED BUILDINGS
MIKAYEL G. MELKUMYAN Earthquake Engineering Center of the National Survey for Seismic Protection under the Government of Armenia, Davidashen -IVMassiv, Yerevan, 375054, Armenia E-mail:
[email protected] In recent years in Armenia the introduction of seismic isolation into the practice of new construction as well as of seismic protection of existing buildings has started. In 1992 EEC of NSSP has launched active works searching for new solutions which could strengthen buildings and structures against the attacks of the underground element. New technologies, using seismic isolation systems for upgrading the earthquake resistance of existing buildings were developed in EEC, which have already attracted international professional attention. They allow upgrading the earthquake resistance of existing buildings without interruption of their functioning. Along with that seismic isolation systems were developed and designed for the construction of different new buildings. The paper presents the state of the art in structural control in Armenia [1, 2, 3, 4] and devoted to the description of the structural concepts, which can provide significant seismic safety for existing and newly constructed buildings.
1
The first attempt to design building with seismic isolation
The study of the experience accumulated in the field of seismic isolation worldwide became the basis for the first attempt to design a seismic isolation system. An experimental (pilot) project of a 4-story R/C frame building with seismic isolation was developed [5]. The number of bearings in the isolation system was 22. Vibration period of the building with base isolation was taken 1.1 sec as it was planed to test this building in the resonance regime using special vibration machine. With the bigger period of the considered building it is impossible to realize the test, as the excitation force on the vibrator shafts will decrease significantly. The design displacement of isolation system was 35 mm. The laminated rubber bearings (LRB) were designed using different approaches [6, 7, 8]. Their comparison shows a substantial disparity of the values of elasticity characteristics. However, taking into account the lack of results of our own experimental studies, the technique given in [7] was chosen. Parameters of the LRBs are presented below in Table 1. After the designing was over EEC together with the "NATRIT" plant manufactured in autumn 1993 the first in Armenia LRB using cold fastening technique. The free vibrations of the LRB were recorded when constant vertical load was applied. The damping of the LRB was equal to about 3% and the period of vibrations - 1.2 sec. Then under the horizontal force the bearing was gradually brought to failure. Force - displacement relationships in the form of hysteresis loops were obtained. The performance of the bearing up to horizontal displacement of 55 365
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mm was elastic. With further increase of horizontal displacement cracks occurred in the lower layers of the bearing due to deterioration of rubber-metal bonding. Test proves that Armenia is able to produce seismic isolation structures and apply them to solve the main task of seismic risk reduction. Table 1. Parameters of laminated rubber bearings
2
Name of parameter
Symbol of unit
Value
Overall height Overall diameter Number of rubber layers Thickness of rubber layer Number of steel layers Thickness of steel layer Diameter of steel layer Thickness of rubber cover layer (side) Thickness of endplates Diameter of two endplates Shear stiffness Vertical load
mm mm
158.7 368.0 19.0 4.8 18.0 1.5 360.0 4.0 20.0 510.0 0.049 220.0
mm mm mm mm mm mm kN/mm kN
First application of seismic isolation in Armenia
In 1995 the first real application of seismic isolation system in the record of earthquake engineering in Armenia was realized on a one-story bathhouse. The project was developed in cooperation with Engineering Research Center of the American University of Armenia. It has been introduced in Giumri, Vanadzor and Spitak. Each of bathhouse rests on 21 seismic isolation bearing [9]. Design displacement of bearings was 10 cm. Totally "NATRIT" plant has manufactured 193 bearings with the parameters listed in Table 2. A question may arise of whether it was reasonable to apply seismic isolation system at a simple one-story building. Our opinion is that to ensure successful evolution of this new earthquake engineering line one should proceed from simple to complicated tasks gaining new experience with each step. On the other hand the design envisions placing 20 ton capacity tanks for cold and hot water in attic space of the bathhouses. This will adversely affect capability of the buildings to withstand reliably seismic impacts. However, the seismic isolation nullifies such a negative effect of water tanks.
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Table 2. Parameters of bearings for bathhouse Name of parameter
Symbol of unit
Value
Overall height Overall diameter Number of rubber layers Thickness of rubber layer Number of steel layers Thickness of steel layer Diameter of steel layer Thickness of rubber cover layer (side) Thickness of endplates Diameter of two endplates Shear stiffness Vertical load
mm mm
206.0 200.0 20.0 7.0 19.0 2.0 190.0 5.0 14.0 330.0 0.15 75.0
mm mm mm mm mm mm kN/mm kN
Four bearings have been tested at EEC. The natural vibration period and damping factor were measured. Vibration period in the horizontal direction of the bearings was equal to 1.13 sec, while logarithmic decrement was 0.31 that corresponds to 5% of critical damping. The LRBs were tested through exposition to reversal static horizontal forces. Two LRBs were tested simultaneously. Vertical load of 150 kN, simulating the weight of the building on two bearings, was applied to the upper frame of the test machine. Test of LRBs under reversal horizontal loads revealed their satisfactory behavior within the range of design displacement values. In total, 4 cycles of deforming were realized. At each following cycle the displacement was increased by 2 cm. Upon achieving a displacement of 8 cm the specimens were deformed in single direction by increasing the horizontal load up to the failure. The force - displacement relationship was obtained. As it demonstrates, approaching the design value of displacements, performance of LRB changes to clearly defined nonlinear stage. At this stage of deforming an increase in bearing stiffness is observed. Failure occurred under 14 cm displacement of the specimens. As the specimen failed, both its glue and rubber layers partly ruptured. Such type of failure is evidence to the fact that for the given rubber type the best strength of rubber-tometal bonding was really achieved. The outcome of the testing appears to be encouraging. Being this time manufactured using "hot fastening" technology, the LRBs have been collapsing only when horizontal deformation appreciably exceeded design values.
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3
Control of seismic forces at upgrading the earthquake resistance of existing 9-story R/C frame building
The 9-story R/C buildings are designed as prefabricated framed systems, the horizontal stiffness of which is provided in the longitudinal direction by the frames with strong beams and in the transverse direction by the frames with weak beams and shear walls [10]. These buildings were heavily damaged during the 1988 Spitak earthquake. The project on upgrading seismic resistance of a 9-story building by means of Additional Isolated Upper Floor (AIUF) method [11] pioneered in applying seismic isolation structures to the top part of the building instead of its base and was implemented in 1995. The considered building has the square plan with the distance between columns 6x6 m. There are 16 columns in the plan of the building. All columns are passing through the slab of the ninth floor on the height of 1.0 m into the space of the attic floor. The assembling of AIUF over the 9-story building starts after dismantling of the attic floor. The connection of AIUF to the building was designed by means of rubber bearings. In this project the high damping rubber bearings (HDRBs) were used. The bearings were designed in collaboration with Tun Abdul Razak Research Center (TARRC) to give a horizontal displacement of 13 cm. All 16 columns on the top of the building were taken into steel jackets with the height of 1.0 m so that the horizontal parts of each jacket in the size of 414x414 mm represented steel plates with the thickness of 25 mm, to which recess rings were bolted. The steel jackets of all 16 columns were connected to each other by means of steel trusses. Thus a rigid structure is created to transfer the forces from AIUF to the building. To implement the project 16 bearings were manufactured by Min Rubber Products Sdn. Bhd., Beranang, Selangor, Malaysia. Parameters of the HDRBs are presented below in Table 3. After installation of all HDRBs the structures of AIUF were assembled above them. The AIUF represents a steel frame structure with the same number of columns in plan as in the building. The base of each column is a steel plate bolted to the upper recess rings of HDRBs. All steel columns of AIUF also were connected to each other by means of steel trusses. On the level of upper belts of trusses a R/C slab is designed using precast panels. In essence, the additional floor itself also represents the rigid structure, which during the earthquakes, being supported by HDRBs, practically has no deformations. Under the earthquake impact AIUF, acting as vibration damper, reduces stressdeformed state of the building and increases its earthquake resistance in average by a factor of 1.6 [11]. The reduction of shear forces horizontal displacements in the building with AIUF takes place because of increase of the period of vibration of the whole system (building plus AIUF). Then a new type of second mode of vibration appears and becomes prevailing and as a result AIUF oscillates in anti-phase related to the building.
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Table 3. Parameters of bearings for AIUF Name of parameter
Symbol of unit
Value
Overall height Overall diameter Number of rubber layers Thickness of rubber layer Number of reinforcing layers Thickness of reinforcing layer Diameter of reinforcing layer Thickness of rubber cover layer (side) Thickness of endplates Diameter of two endplates Thickness of end cover layer Shear stiffness Vertical load
mm mm
196.0 380.0 14.0 9.0 13.0 2.0 360.0 10.0 20.0 376.0 2.0 0.81 820.0
mm mm mm mm mm mm mm kN/mm kN
It is worth noting that the isolated upper floor allows not only upgrading earthquake resistance of a building, but enlarging its useful space as well. The most distinctive feature of the new earthquake resistance upgrading method, however, is that there is no need to re-settle residents from the building during construction works. In present time the upgrading of earthquake resistance of two buildings by means of AIUF is already accomplished in the city of Vanadzor.
The use of high damping rubber isolators for retrofitting of existing 5-story stone building Buildings of this type also have been erected in all regions of Armenia. They have the bearing walls located mainly in transverse direction. The horizontal stiffness in the longitudinal direction is provided partly by the R/C frames with strong beams and columns, made inside the body of walls, and by longitudinal walls at the edge parts of the buildings. The most vulnerable zones in these buildings are the edge parts where the direction of bearing walls had been changed. It is in these very zones that intensive plastic deformations resulting in failure of the buildings have been developed due to the weak connections between longitudinal and transverse wall [12]. The developed by EEC in 1994 structural concept aims to retrofit an existing building by means of seismic isolators using simple working technology. This is a unique pioneering seismic isolation project introduced in 1996 for an existing 5-
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story stone building. The idea is to supply this building with seismic isolation by gradually cutting the building from its foundation. It was supposed to install the isolators at the level of foundation upper edge by creating continuous upper and lower R/C beams along all bearing walls of the building. After removing parts of walls between seismic isolators the building appears separated from its foundation and linked to it only through seismic isolators. It is very important that openings in walls are made with single-spacing, i.e., two adjacent openings should not be made simultaneously; parts of walls existing between seismic isolators should be cut off beginning from the middle of building in plan. The operation is made without resettlement of the dwellers. The world practice has had no similar precedent in retrofitting of apartment buildings. To implement the project two types of bearings were used [13]. The seismic isolators were designed in collaboration with TARRC to give a horizontal displacement of 13 cm. Min Rubber Products Sdn. Bhd., Sime Engineering Rubber Products Sdn. Bhd. and MRPRA have manufactured all together 60 HDRBs (see Table 3) in Malaysia and UK. In present time the retrofitting of one building by means of base isolation is already accomplished also in the city of Vanadzor.
5
Implementation of base isolation in construction of new four story apartment building
In 1996 for the first time in Armenia a seismic isolation system was designed for the construction of a new apartment building in the center of the old Spitak - the city that was destroyed during the Earthquake of December 7, 1988. The structure of the building represents a system with monolithic bearing walls and prefabricated slabs. Thirty-nine HDRBs manufactured in Malaysia by Min Rubber Products Sdn. Bhd. were used for this building (see Table 3). Two unique tests were carried out in 1997, when construction of the building was almost completed [4]. The first one was the trial of the technology of replacement of seismic isolators. During the construction a dummy isolator, made from the steel pipe, was installed in certain positions instead of a rubber bearing. Two jacks with the capacity of lOOt were used to lift the building at each location by about 0.5 mm. After that the dummy isolator was taken out and installation of the real seismic isolator started. The latter was gradually brought into its design position with the recess rings placed around it. Finally the two rings were bolted to the upper and lower steel plates. The trial confirms that replacement of the isolators, should this ever be necessary, can be accomplished in a quick and simple operation. In order to carry out the second test, a special loading system by which a horizontal static or dynamic force could be applied to the building at the level of the upper beams of the isolation system was designed and constructed near the building. Four cycles of loading and unloading were performed during the static tests up to a maximum displacement of 20% of the design value. It was revealed that the initial
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stiffness (at 0.2 cm displacement) of the isolation system is more than 10 times higher than its expected stiffness at the design displacement; the ability of the system to provide an intrinsic restraint against wind loading is confirmed. The nonlinear behavior of the isolation system is close to that observed in tests on rubber samples. At a deformation equal to about 20% of the design value the observed secant stiffness of the overall isolation system correlates well with the stiffness expected at that deformation from the design calculations and the quasistatic test results on individual isolators [13]. The dynamic tests were carried out by releasing the isolation system at the displacement of 0.2 cm. The experimental value of the period of vibrations is virtually identical with the value calculated on the basis of initial stiffness of the isolation system and actual weight of the building. The displacements and accelerations at the third floor and the roof are respectively 1.03 and 1.06 times those at the first floor, figures agreeing well with the design analysis. The damping of the isolation system is equal to 8.8% of critical damping at small displacements (
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