Damping technologies for tall buildings : theory, design guidance and case studies 9780128159644, 0128159642

Citation preview

Damping Technologies for Tall Buildings Theory, Design Guidance and Case Studies

This page intentionally left blank

Damping Technologies for Tall Buildings Theory, Design Guidance and Case Studies

Alberto Lago Research Assistant, CTBUH/Iuav University of Venice, Venice, Italy

Dario Trabucco Research Manager, CTBUH/Iuav University of Venice, Venice, Italy

Antony Wood Executive Director, CTBUH, Chicago, United States

Butterworth-Heinemann is an imprint of Elsevier The Boulevard, Langford Lane, Kidlington, Oxford OX5 1GB, United Kingdom 50 Hampshire Street, 5th Floor, Cambridge, MA 02139, United States Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved. No part of this publication may be reproduced or transmitted in any form or by any means, electronic or mechanical, including photocopying, recording, or any information storage and retrieval system, without permission in writing from the publisher. Details on how to seek permission, further information about the Publisher’s permissions policies and our arrangements with organizations such as the Copyright Clearance Center and the Copyright Licensing Agency, can be found at our website: www.elsevier.com/permissions. This book and the individual contributions contained in it are protected under copyright by the Publisher (other than as may be noted herein). Notices Knowledge and best practice in this field are constantly changing. As new research and experience broaden our understanding, changes in research methods, professional practices, or medical treatment may become necessary. Practitioners and researchers must always rely on their own experience and knowledge in evaluating and using any information, methods, compounds, or experiments described herein. In using such information or methods they should be mindful of their own safety and the safety of others, including parties for whom they have a professional responsibility. To the fullest extent of the law, neither the Publisher nor the authors, contributors, or editors, assume any liability for any injury and/or damage to persons or property as a matter of products liability, negligence or otherwise, or from any use or operation of any methods, products, instructions, or ideas contained in the material herein. British Library Cataloguing-in-Publication Data A catalogue record for this book is available from the British Library Library of Congress Cataloging-in-Publication Data A catalog record for this book is available from the Library of Congress ISBN: 978-0-12-815963-7 For Information on all Butterworth-Heinemann publications visit our website at https://www.elsevier.com/books-and-journals

Publisher: Matthew Deans Acquisition Editor: Ken McCombs Editorial Project Manager: Andrae Akeh Production Project Manager: Kamesh Ramajogi Cover Designer: Christian J. Bilbow Front Cover Image: Shanghai Tower Damper © Connie Zhou Photography | Building Design by Gensler Typeset by MPS Limited, Chennai, India

Council on Tall Buildings and Urban Habitat The Council on Tall Buildings and Urban Habitat (CTBUH) is the world’s leading resource for professionals focused on the inception, design, construction, and operation of tall buildings and future cities. Founded in 1969 and headquartered at Chicago’s historic Monroe Building, the CTBUH is a not-for-profit organization with an Asia Headquarters office at Tongji University, Shanghai; a research office at Iuav University, Venice, Italy; and an academic office at the Illinois Institute of Technology, Chicago. CTBUH facilitates the exchange of the latest knowledge available on tall buildings around the world through publications, research, events, working groups, web resources, and its extensive network of international representatives. The council’s research department is spearheading the investigation of the next generation of tall buildings by aiding original research on sustainability and key development issues. The council’s free database on tall buildings, The Skyscraper Center, is updated daily with detailed information, images, data, and news. The CTBUH also developed the international standards for measuring tall building height and is recognized as the arbiter for bestowing such designations as “The World’s Tallest Building.” CTBUH Headquarters The Monroe Building 104 South Michigan Avenue, Suite 620 Chicago, IL 60603, United States Phone: 11 (312) 283
5599 Email: [email protected] www.ctbuh.org www.skyscrapercenter.com CTBUH Asia Headquarters College of Architecture and Urban Planning (CAUP) Tongji University 1239 Si Ping Road, Yangpu District Shanghai 200092, China Phone: 1 86 21 65982972 Email: [email protected]

v

vi

Council on Tall Buildings and Urban Habitat

CTBUH Research Office Iuav University of Venice Dorsoduro 2006 30123 Venice, Italy Phone: 139 041 257 1276 Email: [email protected] CTBUH Academic Office S.R. Crown Hall Illinois Institute of Technology 3360 South State Street Chicago, IL 60616, United States Phone: 11 (312) 567 3487 Email: [email protected]

“The new generation of high rise buildings are demanding higher specifications. Damping technologies need to be adapted to meet these needs in a changing environment. Bouygues as a market leader in construction have funded the research programme with CTBUH to develop a guide on to how to work best with these technologies and to improve the conditions of tall buildings for the future”. Marc BLONDEAU, Fabienne FOUCAULT and Andre´ LY – Structure Design Department of Bouygues Batiment International.

Philippe DEFOSSEZ—Technical Director
Head of Central Engineering Department of Bouygues Batiment International. Alain BACHELIER—Technical Director of Central Engineering Department of Bouygues Batiment International. Christian CREMONA—Technical Director of R&D Innovation of Bouygues Construction. Christian de NACQUARD—Technical Director of R&D Innovation of Bouygues Batiment International. Magdalena PYSZKOWSKA—Head of BIM Deployment of Bouygues Construction.

vii

This page intentionally left blank

Contents Acknowledgments ...................................................................................................xv

CHAPTER 1 Executive summary ....................................................... 1 CHAPTER 2 Introduction ................................................................... 7 2.1 Aim of the Book...........................................................................10 2.2 Terms and Definitions..................................................................12 2.3 List of Symbols and Abbreviations .............................................21 2.3.1 Abbreviations .................................................................... 21 2.3.2 Symbols ............................................................................. 21 2.4 History of Dynamic Modification Devices in Tall Buildings.....28 2.5 Industrial Application of Damping Devices................................30 2.5.1 Vibration Control .............................................................. 30 2.5.2 Vibration and Seismic Control of Industrial Machine Structures ........................................................... 32 2.5.3 Vibration Isolation of Buildings and Rail Structures....... 34 2.5.4 Vibration Control of Large Industrial Chimney Stacks ................................................................................ 34

CHAPTER 3 Damping considerations in tall buildings.................. 39 3.1 Basic Principles of Building Response........................................41 3.1.1 Single Degree-of-Freedom System................................... 41 3.1.2 Multiple Degrees-of-Freedom System.............................. 47 3.1.3 Equivalent Viscous Damping ........................................... 53 3.2 Damping in Tall Buildings ..........................................................55 3.2.1 Damping Modeling ........................................................... 55 3.2.2 Intrinsic Damping ............................................................. 57 3.3 Effect of Damping on Building ...................................................73 3.3.1 Other Development Considerations.................................. 77 3.4 Tall Buildings Wind-Excited Motion ..........................................78 3.4.1 Building Wind Vibration .................................................. 79 3.4.2 Occupant Comfort............................................................. 80 3.4.3 Wind Deflection Criteria .................................................. 84 3.5 Tall Buildings Earthquake-Excited Motion.................................85 3.5.1 Principles of Performance-Based Earthquake-Resistant Design for Tall Buildings ................................................. 87 3.5.2 The Role of Damping in the Seismic Response Control of Tall Buildings.................................................. 89 3.5.3 Seismic Deflection Criteria............................................... 90

ix

x

Contents

3.6 Environmental and Economic Considerations.............................90 3.6.1 Life-Cycle Assessment ..................................................... 93 3.6.2 Cost Implication of Buildings with Dissipation Devices ............................................................................ 101 3.7 Damping Technology Uncertainty and Robustness Performance................................................................................102 3.8 Alternative to Damping Devices................................................103

CHAPTER 4 An introduction to dynamic modification devices... 107 4.1 Passive Damping Systems..........................................................108 4.1.1 Distributed Damping Approaches .................................. 109 4.1.2 Mass Damping Approaches ............................................ 195 4.2 Seismic Isolation ........................................................................207 4.2.1 Base-Isolation Types....................................................... 210 4.2.2 Base-Isolation Manufactures .......................................... 215 4.3 Active, Semiactive, and Hybrid Systems ..................................217 4.3.1 Active Systems................................................................ 222 4.3.2 Hybrid Systems ............................................................... 224 4.3.3 Semiactive Systems ........................................................ 226 4.3.4 Adaptive Tuned-Mass Damper Systems ........................ 229 4.3.5 Control Strategies............................................................ 230 4.3.6 Future Directions............................................................. 231 4.3.7 Active, Semiactive, and Hybrid Dampers Manufactures................................................................... 232 4.4 Comparison of Dampers in Tall Buildings................................232

CHAPTER 5 Design procedures for tall buildings with dynamic modification devices ................................. 235 5.1 Available Codes and Design Tools ...........................................236 5.1.1 Codes and Guidelines ..................................................... 236 5.1.2 Practical Design Aspects ................................................ 248 5.1.3 Structural Analyses ......................................................... 251 5.2 Passive Damping Systems..........................................................264 5.2.1 Step-by-Step Procedure for Distributed Dampers.......... 265 5.2.2 Step-by-Step Procedure for Mass Dampers ................... 321 5.3 Isolation Systems........................................................................347 5.3.1 Step-by-Step Procedure for Base Isolation .................... 347 5.4 Active, Semiactive, and Hybrid Systems ..................................365 5.4.1 Literature Review............................................................ 365 5.4.2 Step-by-Step Procedure................................................... 367 5.5 Retrofit of Existing Buildings....................................................400 5.5.1 Code Requirements ......................................................... 400

Contents

5.5.2 Evaluation Procedures Based on ASCE 41-13 (ASCE, 2013).................................................................. 402 5.5.3 Step-by-Step Procedure................................................... 403 5.5.4 Case Study: Retrofitting Examples of High-Rise Building with Damping Systems .................................... 422 5.6 Dynamic Modification Devices Strategy Optimization ............429 5.6.1 Introduction ..................................................................... 429 5.6.2 Algorithm-Based Optimization Procedures.................... 430 5.6.3 Nonalgorithm-Based Optimization Procedures .............. 433

CHAPTER 6 Architectural aspects and building system interaction ................................................................. 437 6.1 Architectural Aspects .................................................................438 6.1.1 Distributed Damping Systems ........................................ 439 6.1.2 Mass Damping Systems.................................................. 444 6.1.3 Base Isolation Systems ................................................... 447 6.2 Elevators .....................................................................................450 6.3 Mechanical Systems ...................................................................455 6.3.1 Basic Considerations....................................................... 455 6.3.2 Mechanical Floors........................................................... 456 6.4 Fac¸ade.........................................................................................457 6.4.1 Double-Skin Fac¸ade........................................................ 458 6.4.2 Diagrid Fac¸ade ................................................................ 460 6.4.3 Mega Brace Dampers Fac¸ade ......................................... 463

CHAPTER 7 Testing, inspection, and maintenance..................... 465 7.1 Codes and Standards Development ...........................................466 7.1.1 US Standards................................................................... 466 7.1.2 Japanese Code ................................................................. 470 7.1.3 Chinese Code .................................................................. 471 7.1.4 European Standards......................................................... 472 7.2 Preinstallation Tests and Quality Control..................................472 7.2.1 Distributed Damping Systems ........................................ 473 7.2.2 Mass, Active, Semiactive, and Hybrid Damping Systems ........................................................... 482 7.2.3 Base Isolation Systems ................................................... 486 7.2.4 Testing Examples............................................................ 491 7.2.5 Summary ......................................................................... 493 7.3 Commissioning and System Tuning ..........................................493 7.3.1 Distributed Damping Systems ........................................ 494 7.3.2 Mass Damping Systems.................................................. 495 7.3.3 Base Isolation Systems ................................................... 500

xi

xii

Contents

7.4 Fatigue of Dampers ....................................................................501 7.4.1 Viscous Dampers ............................................................ 502 7.4.2 Viscoelastic Dampers...................................................... 503 7.4.3 Displacement-Dependent Dampers ................................ 504 7.4.4 Mass Damping Systems.................................................. 504 7.4.5 Base Isolation Systems ................................................... 505 7.5 Building Health Monitoring.......................................................505 7.5.1 Distributed Damping Systems ........................................ 507 7.5.2 Mass Damping Systems.................................................. 507 7.5.3 Active, Semiactive, and Hybrid Damping Systems ....... 507 7.5.4 Base Isolation Systems ................................................... 508 7.5.5 Case Study Examples...................................................... 508 7.6 Ongoing Maintenance ................................................................513 7.6.1 Standard Recommendations............................................ 515 7.6.2 Distributed Damping Systems ........................................ 516 7.6.3 Mass Damping Systems.................................................. 516 7.6.4 Active, Semiactive, and Hybrid Damping Systems ....... 517 7.6.5 Base Isolation Systems ................................................... 518 7.7 Maintenance Cost.......................................................................519 7.8 Postextreme Event Inspections ..................................................519 7.8.1 Distributed Damping Systems ........................................ 521 7.8.2 Mass Damping Systems.................................................. 522 7.8.3 Base Isolation Systems ................................................... 523 7.9 Transportation, Installation, and Care of Works .......................524 7.9.1 Transportation of Damping Systems .............................. 524 7.9.2 Storage/Installation of Damping Systems ...................... 525 7.9.3 Care of Works of Damping Systems.............................. 526 7.10 Resilience-Based Earthquake Design Initiative and United States Resiliency Council Rating Systems .............526 7.10.1 Resilience-Based Earthquake Design Initiative Rating System ............................................................... 527 7.10.2 United States Resiliency Council Rating System ........ 529

CHAPTER 8 Case studies of tall buildings with dynamic modification devices ................................................ 533 8.1 Distributed Damping Systems Case Studies .............................534 8.1.1 Columbia Tower, Seattle, Washington, United States ................................................................. 534 8.1.2 Two Union Square, Seattle, Washington, United States ............................................................................. 547 8.1.3 St Francis Shangri La Place, Manila, Philippines........ 556

Contents

8.1.4 Pangu Plaza, Beijing, China ......................................... 567 8.1.5 Beijing Yintai Center, Beijing, China .......................... 577 8.1.6 San Diego Central Courthouse, San Diego, California, United States............................................... 592 8.1.7 Wuhan Poly Cultural Plaza, Wuhan, China ................. 615 8.1.8 TianJin International Trade Center, TianJin, China ............................................................................. 626 8.1.9 454 Yonge, Toronto, ON, Canada................................ 637 8.1.10 181 Fremont Street, San Francisco, California, United States ................................................................. 658 8.1.11 Atushi Building, Xin Jiang, China ............................... 668 8.1.12 Costums Residential, Auckland, New Zealand ............ 678 8.1.13 Connor Tower, Manila, Philippines ............................. 701 8.1.14 Allianz Tower, Milan, Italy .......................................... 716 8.2 Mass Damping Systems Case Studies .......................................733 8.2.1 Citicorp Building, New York, New York City, United States ................................................................. 734 8.2.2 John Hancock Tower, Boston, Massachusetts, United States ................................................................. 745 8.2.3 Taipei 101, Taipei, Taiwan........................................... 758 8.2.4 One Rincon Hill (South Tower) San Francisco, California, United States............................................... 774 8.2.5 Comcast Center, Philadelphia, Pennsylvania, United States ................................................................. 783 8.2.6 Hyatt Park Tower, Chicago, Illinois, United States..... 788 8.2.7 Highcliff Apartments, Hong Kong ............................... 801 8.2.8 Bloomberg Tower, New York City, New York, United States ................................................................. 808 8.2.9 Raffles City, Chongqing, China ................................... 818 8.2.10 L-Tower, Toronto, Canada ........................................... 828 8.2.11 14 York Street, Toronto, Canada.................................. 840 8.2.12 One Bloor Street East, Toronto, Canada ...................... 847 8.2.13 1151 West Georgia, Vancouver, Canada ..................... 854 8.2.14 Shanghai Tower, Shanghai, China ............................... 863 8.2.15 The Independent, Austin, Texas, United States ........... 876 8.3 Base Isolation Systems Case Studies.........................................888 8.3.1 Nunoa Capital Building, Santiago, Chile ....................... 888 8.4 Active, Semiactive, and Hybrid Systems Case Studies ............904 8.4.1 Thyssenkrup Test Tower, Rottweil, Germany ............... 904

xiii

xiv

Contents

CHAPTER 9 Future of dynamic modification systems ................. 921 9.1 Improved Modeling of Structural Behavior ..............................922 9.1.1 Improved Understanding of Actions on Structures........ 922 9.1.2 Improved Behavior Model of Materials and Devices.... 923 9.1.3 Enhanced Computing Power .......................................... 923 9.1.4 Improved Design Codes.................................................. 924 9.2 Implementing Results of Technological Progress .....................925 9.2.1 Development of Existing Devices and New Hybrid Systems............................................................................ 925 9.2.2 New Technologies........................................................... 926

CHAPTER 10 Tall building with dynamic modification systems trend data.................................................... 929 10.1 Database 1 (Worldwide Buildings Over 250 m).......................930 10.1.1 General Trends for Tall Buildings................................ 930 10.1.2 General Trends for Tall Buildings with Dynamic Modification Systems ................................................... 930 10.1.3 Trends for Dynamic Modification System Category ........................................................................ 937 10.1.4 Trends for Dynamic Modification System Types........ 941 10.2 Database 2 (US Buildings Over 200 m)....................................943 10.2.1 General Trends for Tall Buildings................................ 943 10.2.2 General Trends for Tall Buildings with Dynamic Modification System ..................................................... 945 10.2.3 Trends for Dynamic Modification System Category ........................................................................ 947 10.2.4 Trends for Dynamic Modification System Type ......... 948 10.3 Further Studies ...........................................................................948 10.3.1 Dynamic Modification System Versus Height............. 949 10.3.2 Structural Material ........................................................ 953 10.3.3 Structural System .......................................................... 958 10.3.4 Building Function ......................................................... 958 10.4 Summarized Data for Tall Buildings with Dynamic Modification System ..................................................................961 10.5 Conclusions ................................................................................961

CHAPTER 11 Conclusions............................................................... 969 Appendix A............................................................................................................971 References............................................................................................................1013 Index ....................................................................................................................1067

Acknowledgments The Council on Tall Buildings and Urban Habitat (CTBUH) acknowledges the work of several international experts who have contributed different chapters. This group comprises individuals from many backgrounds, including consulting engineering, research, construction industry, education, and developers. The individuals who contributed for different chapters as main authors, contributors, main and general peer reviewers are listed below.

MAIN AUTHORS Chapter 1 Chapter 2 Chapter 3

Chapter 4

Chapter 5

Chapter 6

Chapter 7 Chapter 8

Gary Hart (Thorton Tomasetti), Dario Trabucco (CTBUH) Alberto Lago and Dario Trabucco (CTBUH), Christian Meinhardt (GERB Vibration Control Systems) Donald Davies and Farshad Berahman (MKA), Rob Smith (Arup), Stefano Cammelli (BMT), Mario Lafontaine (Rene Lagos Engineers), Yongqi Chen (Beijing QITAI), Alberto Lago (CTBUH) Alan Klembczyk, Craig W. Winters and David Lee (Taylor Devices), Yongqi Chen and Cheng Peng (Beijing QITAI), Bogdan Catalin Stefanescu (UTCB, Bucharest), Christian Meinhardt (GERB Vibration Control Systems), Mike Tait (McMaster Unviersity), Michael Montgomery (Kinetica Dynamics), Alberto Lago and Hadi Moghadasi Faridani (CTBUH) Alberto Lago and Hadi Moghadasi Faridani (CTBUH), Robert McNamara and Bart Sullivan (McNamara Salvia), Kevin MacLean and Tibor Kokai (Read Jones Christoffersen Ltd.), Mario Lafontaine (Rene Lagos Engineers) Peter Weismantle and Sara Beardsley (Smith Gill), Stefan Kaczmarcyk (University of North Hampton), Hadi Moghadasi Faridani and Alberto Lago (CTBUH) Peter Lee (SOM), Elena Mola and Carlo Segato (ECSD), Alan Klembczyk (Taylor Devices), Hadi Moghadasi Faridani and Alberto Lago (CTBUH) 8.1.1 Donald Davies and Farshad Berahman (MKA) 8.1.2 Donald Davies and Farshad Berahman (MKA) 8.1.3 Rob Smith (Arup) 8.1.4 Yongqi Chen (Beijing QITAI) 8.1.5 Yongqi Chen (Beijing QITAI) 8.1.6 Peter Lee (SOM) 8.1.7 Yongqi Chen (Beijing QITAI) 8.1.8 Yongqi Chen (Beijing QITAI), Aaron Wang (Capital Land) 8.1.9 Michael Montgomery (Kinetica Dynamics) 8.1.10 Rob Smith (Arup) 8.1.11 Yongqi Chen (Beijing QITAI) 8.1.12 Nick Gillespie (Mott MacDonald), Ignacio Vial and Michael Rendel (Sirve) (Continued)

xv

xvi

Acknowledgments

Continued 8.1.13

Chapter 9 Chapter 10 Chapter 11 Appendix A

Michael Montgomery (Kinetica Dynamics), Jose Sy (Sysquared), Gerald Delos Reyes 8.1.14 Elena Mola (ECSD) and Carlo Segato 8.2.1 Craig D. Blanchet (Le Messurier) 8.2.2 Craig D. Blanchet (Le Messurier) 8.2.3 Derek Kelly, Jon Galsworthy, Tanya Martinez (RWDI), Trevor Haskett and Andy Smith (Motioneering), Dennis Poon (Thornton Tomasetti) 8.2.4 Donald Davies and Farshad Berahman (MKA) 8.2.5 Derek Kelly, Jon Galsworthy, Tanya Martinez (RWDI), Trevor Haskett and Andy Smith (Motioneering), Dennis Poon (Thornton Tomasetti) 8.2.6 Derek Kelly, Jon Galsworthy, Tanya Martinez (RWDI), Trevor Haskett and Andy Smith (Motioneering), Dennis Poon (Thornton Tomasetti) 8.2.7 Donald Davies and Farshad Berahman (MKA) 8.2.8 Derek Kelly, Jon Galsworthy, Tanya Martinez (RWDI), Trevor Haskett and Andy Smith (Motioneering), Dennis Poon (Thornton Tomasetti) 8.2.9 Li-Gang Zhu (Arup), Yongqi Chen (Beijing QITAI), Aaron Wang (Capital Land) 8.2.10 Un Yong Jeong (Gradient Wind) 8.2.11 Un Yong Jeong (Gradient Wind) 8.2.12 Un Yong Jeong (Gradient Wind) 8.2.13 Un Yong Jeong (Gradient Wind) 8.2.14 Derek Kelly, Jon Galsworthy, Tanya Martinez (RWDI), Trevor Haskett and Andy Smith (Motioneering), Dennis Poon (Thornton Tomasetti) 8.2.15 Tony Rofail (WindTech) 8.3.1 Mario Lafontaine (Rene Lagos), Rodrigo Retamales (Rubén Boroschek & Asociados) 8.4.1 Christian Meinhardt (GERB Vibration Control Systems) Bogdan Catalin Stefanescu (UTCB, Bucharest) Hadi Moghadasi Faridani, Alberto Lago and Dario Trabucco (CTBUH) Alberto Lago and Dario Trabucco (CTBUH) Alan Klembczyk, Craig W. Winters and David Lee (Taylor Devices), Alberto Lago and Hadi Moghadasi Faridani (CTBUH), Yongqi Chen and Cheng Peng (Beijing QITAI), Bogdan Catalin Stefanescu (UTCB, Bucharest), Christian Meinhardt (GERB Vibration Control Systems), Mike Tait (McMaster Unviersity), Michael Montgomery (Kinetica Dynamics)

Acknowledgments

CONTRIBUTORS Chapter 3 Chapter 4

Chapter 5 Chapter 7 Chapter 9

Chapter 10 Appendix A

Abdelrahman Elbakeit (Kind Saud University), Amir S.J. Gilani (Miyamoto International, Inc.), Aaron Mazeika (SOM), Aaron Wang (Capital Land) Michael Constantinou (SUNY University of Buffalo), Brial Lim (VSL), David Auclair (Bouygues Constructions), Tony Rofail and Matthew Vallis (WindTech), Kurt Strobel (GERB Vibration Control Systems), Imad Mualla (Damptech) Elena Mola and Carlo Segato (ECSD), Brian Lim (VSL), Jon Galsworthy (RWDI) C.T. Fung (VSL), Aaron Wang (Capital Land) Alberto Lago (CTBUH), Alan Klembczyk and Yongqi Chen (Taylor Devices), Stefano Cammelli (BMT), Michael Montgomery (Kinetica Dynamics), Tony Rofail (Wind Tech), Gianpaolo Colato and Gabriella Castellano (FIP), Christian Meinhardt (GERB Vibration Control Systems). Dennis Poon and Jianhai Liang (Thorton Tomasetti) Michael Constantinou (SUNY University of Buffalo), Brial Lim (VSL), David Auclair (Bouygues Constructions), Tony Rofail and Matthew Vallis (WindTech), Kurt Strobel (GERB Vibration Control Systems), Imad Mualla (Damptech)

MAIN PEER REVIEW PANEL Chapter 2

Gary Hart (Thorton Tomasetti), Ying Zou (Tonji University), Christopher Rojahn (AT Countil), Rodolfo Saragoni (Universidad De Chile), Yozo Fujino (University of Tokio) Chapter 3 Chritian Cruz (Nishkian Menninger), Andrew Taylor (KPFF), Martin Button (Button Engineering), Michael Constatinou (SUNY University of Buffalo), Stephen Main, Jiun-Wei Lai and Shanshan Wang (UC Berkeley) Chapter 4 Donald Davies (MKA), Michale Montgomery (Kinetica Dynamics), Gabriella Castellano (FIP), Tony Tschanz (Sunpad) Chapter 5 Neville Mathias and SOM Structural Office in San Francisco (SOM), Eric Ko (Arup), Rafel Sabelli (Walter P. Moore), Ali Roufegarinejad (Forell/Elsesser Engineers, Inc.), Chung-Soo Doo (BART) Chapter 6 Luca Truggestad (SOM), Mark Walsh (Perkins Will), Robert Sinn (Thornton Tomasetti) Chapter 7 Aaron Wang (Capital Land), Xin Zhao (Tonji University), L.G. Zhu (Arup), Guo Biao Lou (Tonji University) Appendix A Donald Davies (MKA), Michale Montgomery (Kinetica Dynamics), Gabriella Castellano (FIP)

xvii

xviii

Acknowledgments

GENERAL PEER REVIEW PANEL Chapter 2

Amir Gilani (Miyamoto International, Inc.), Bogdan Catalin Stefanescu (UTCB, Bucharest), Carlo Segato (ECSD), Stefano Cammelli (BMT), Christian Meinhardt and Kurt Strobel (GERB Vibration Control Systems), Donald Davies and Farshad Berahman (MKA), Trevor Haskett, Shaney Love and Andy Smith (Motioneering), Kirk Harman (The Harman Group, Inc.), William Douglas Miranda and Alberto Lago(CTBUH), Roy Denoon (CPP Wind) Chapter 3 Amir Gilani (Miyamoto International, Inc.), Bogdan Catalin Stefanescu (UTCB, Bucharest), Carlo Segato (ECSD), Stefano Cammelli (BMT), Donald Davies and Farhsad Berahman (MKA), Trevor Haskett, Shaney Love and Andy Smith (Motioneering), Kirk Harman (The Harman Group, Inc.), Andrè Ly (Bouygues Construction), Aaron Wang (Capital Land), William Douglas Miranda and Alberto Lago (CTBUH), Roy Denoon (CPP Wind), Kurt Strobel (GERB Vibration Control Systems) Chapter 4 Ronald Mayes (Simpson Gumpertz & Heger), Bogdan Catalin Stefanescu (UTCB, Bucharest), Carlo Segato (ECSD), Stefano Cammelli (BMT), Christian Meinhardt and Kurt Strobel (GERB Vibration Control Systems), Farshad Berahman (MKA), Trevor Haskett, Shaney Love and Andy Smith (Motioneering), Andrè Ly (Bouygues Construction), Ahmad Ahmad Rahim and Elfahal Motaz (WSP), Alan Klembczyk (Taylor Devices), Aaron Wang (Capital Land), William Douglas Miranda and Alberto Lago (CTBUH), Un Yong Jeong (Gradient Wind), Amarnath Kasalanati (PEER), Eriksen Conrad and Mohammed Mohammes (DIS Inc) Chapter 5 Stefano Cammelli (BMT), Aaron Wang (Capital Land), Yongqi Chen (Beijing QITAI), Carlo Segato (ECSD), Mario Lafontaine (Rene Lagos Engineers), Baiping Dong (Leisle E. Robertson), Bogdan Catalin Stefanescu (UTCB, Bucharest), Andrè Ly (Bouygues Construction), Alan Klembczyk (Taylor Devices), William Douglas Miranda and Alberto Lago (CTBUH), Dennis Poon, Paul Fu, Robert Sinn and John Peronto (Thornton Tomasetti) Chapter 6 Aaron Wang (Capial Land), Yongqi Chen (Beijing QITAI), Carlo Segato (ECSD), Baiping Dong (Leisle E. Robertson), Bogdan Catalin Stefanescu (UTCB, Bucharest), Andrè Ly (Bouygues Construction), William Douglas Miranda and Alberto Lago (CTBUH), Dennis Poon, Paul Fu, John Peronto (Thornton Tomasetti) Chapter 7 Stefano Cammelli (BMT), Aaron Wang (Capital Land), Yongqi Chen (Beijing QITAI), Carlo Segato (ECSD), Biaping Dong (Leisle E. Robertson), Bogdan Catalin Stefanescu (UTCB, Bucharest), Andrè Ly (Bouygues Construction), Michale Montgomery (Kinetica Dynamics), Alan Klembczyk (Tyalor Devices), Peter Lee (SOM), Alberto Lago (CTBUH), Roy Denoon (CPP Wind), Kurt Strobel (GERB Vibration Control Systems), Dennis Poon, Paul Fu, Robert Sinn and John Peronto (Thornton Tomasetti), Evan Reis (USRC) Appendix A Ronald Mayes (Simpson Gumpertz & Heger), Bogdan Catalin Stefanescu (UTCB, Bucharest), Carlo Segato (ECSD), Stefano Cammelli (BMT), Christian Meinhardt (GERB Vibration Control Systems), Farshad Berahman (MKA), Trevor Haskett, Shaney Love and Andy Smith (Motioneering), Andrè Ly (Bouygues Construction), Ahmad Rahim and Elfahal Motaz (WSP), Alan Klembczyk (Taylor Devices), Aaron Wang (Capital Land), William Douglas Miranda and Alberto Lago (CTBUH)

CHAPTER

Executive summary

1

Dynamic modification devices in tall buildings are an excellent example of how the artistic vision of the architect can be combined with the scientific knowledge of structural engineering to create very special buildings. The structural engineering area of wind/seismic engineering and the experience of dynamic modification devices in tall buildings do not arrive overnight but result from many government, institutional, and private funding sources to take innovative ideas and then the transfer of this technology into practice. This path from idea to technology transfer has been going on for over approximately 50 years with outstanding international research, communication, and a passionate commitment by structural/wind engineer’s communication to develop a better understanding of the wind forces on and the performance of tall buildings when subjected to wind forces. In structural engineering, we call the design of tall buildings using the results of this commitment performance-based design. What is performance-based design? An illustrative example often used to answer this question is to use a flag pole example as shown in Fig. 1.1. When the flag pole is pulled, the pole is subjected to a force and it deforms. Structural engineers study this force versus deformation by first plotting it and then fitting equations to best represent this curve. These equations result from a study of force versus deformation curves from wind and earthquake forces. The end work product is the selection of what is called a structural system which is an assembly of beams, columns, and dampers that control the performance of the building to client-defined acceptable performance goals. For example, as shown in Fig. 1.1, the client may want us to design the tall building to a performance goal to not have a softening of the structure to the wind forces we expect to occur once every 30 years. What learning has been done to apply to developing a performance-based design? This learning can be called technology transfer. Technology transfer comes from contributions around the world that has resulted in realization of the benefits of using dynamic modification devices in tall buildings through a multistep journey that has taken almost a half century. How is this technology transfer applied to developing a performance-based design? This area of technology transfer is called structural reliability. In the theory of structural reliability, tall building demands are the wind/seismic forces and motions imposed on our structural member and system. The ability of the structural system to resist these demands is the capacity of the tall building. To Damping Technologies for Tall Buildings. DOI: https://doi.org/10.1016/B978-0-12-815963-7.00001-4 Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved.

1

2

CHAPTER 1 Executive summary

FIGURE 1.1 The foundation of performance-based design: the understanding and control of applied force versus resultant deformation.

FIGURE 1.2 Performance-based design limit states.

illustrate this language of structural reliability, consider point A in Fig. 1.2 where the curve represents the force versus deformation curve of an actual tall building when subjected to ever-increasing wind/seismic forces. To resist this demand, the steel and concrete in the tall building must experience elongation of the steel and cracking of the concrete. If point A corresponds to a very small force, the steel elongation and microcracking of the concrete corresponds to damage that may not yet visible to the naked eye. But still corresponding to point A, the building’s concrete structural members have absorbed the energy imparted from the wind force and therefore used up some of its total energy capacity.

Executive Summary

With the increase in the wind/seismic force demands, building performance evolves through many different states. For example, as more wind/seismic force is applied, the structure goes from point A to point B on the force versus deformation curve. Point B can, for example, correspond to when the building occupants start experiencing undesirable motions from the wind/seismic forces on the building. Then with greater forces, the system reaches point C which can correspond to the start of the falling off (spalling) of the concrete from the structural members. Then to point D which corresponds to the start of permanent elongation of the steel reinforced bars in the concrete beam or column structural members. Then to point E which is the maximum force the structural system can be designed to resist wind/seismic forces. Then to point F which corresponds to the point of incipit collapse of the entire structural system if subjected to even slightly larger wind/seismic-induced deformations. In the language of structural reliability, all these different points on the force versus deformation curve are defined as the building’s limit states. These limit states are the cornerstones of performance-based design. When the very rapid technology transfer of structural reliability started in the 1950s that resulted in a major redirection of building codes, performance-based design was named limit state design. These limit states and relative mathematical relationships are defined usually from testing where instrumented specimens are tested to define the different limit states. We know from our use of wind turbines to capture and use wind energy, so how do dynamic modification devices in buildings relate to the wind energy imposed on a tall building? Dynamic modification devices are structural members in our tall building’s part of the lateral force resisting system along with other structural members constructed of steel and concrete. Together these structural members absorb the energy of the wind forces and reduce the magnitude of the deformations that other structural members experience. This can be viewed in our example just discussed, the dynamic modification devices reduce the response such that for a given magnitude of wind force, the building responds in a prior and less undesirable limit state. The benefit of energy absorption devices can be explained with the example of the water from a severe rainstorm going into a mountain lake. If the water is too much in the lake, we remove the water in a confident and predictable way. In a building, we use the dynamic modification devices to remove the energy that a severe windstorm impacts on our buildings’ structural system (i.e., beams, columns, walls, and dampers) in a very confident and predictable way. Concrete structures without dynamic modification devices “eat up” energy with their aggregate interaction and this is acceptable but dynamic modification devices, because they are manufactured under superior quality control, not only eat up the energy but very importantly do so in a more confident and predictable way. How do we use these equations to estimate building performance? The very foundation of structural engineering has always been to use our education and experience at a minimum to develop structural designs that prevent collapse. Especially in the early 20th century, we augmented this education and experience

3

4

CHAPTER 1 Executive summary

with the focus of our learning being the development of equations to estimate structural member and system performance. This resulted, especially starting in the early 1960s, with structural engineers placing a priority on the development and use of laboratory testing results to develop with the benefit of the digital computer a way to not only write equations but also do so by visualizing the structure as a composition of discreet small microstructures that can be called “finite elements.” Each of these finite elements is a microstructure whose force versus deformation performance is represented by equations based on validated laboratory testing results. The determination of the wind/seismic force demand on a structural member or structural system limit state is required by the structural engineer when a structural reliability analysis is performed. For example, major advances in wind engineering have focused on the determination of the wind force demand in the last 50 years since the first world wind engineering conference. Advances can be thought of in two basic areas. One area is the determination of the wind speeds in multiple directions for a specific building site. The results of the advances in this area can not only be obtained for journal papers and conference proceedings but also on the television/web. The other area of advances is the determination of the wind forces on the building in a local surface area or the building in total. The other area is the advances in wind tunnel testing and its companion computational fluid mechanics. Wind tunnel models, which incorporate the natural and built environment, are used to obtain pressure taps on the scale model of the tall building, the pressures on the building surface, and the forces that are input to structural analysis models. The advances in demand enable to obtain best estimates of the expected wind force demand on a limit state that defines performance as well as numerical values of uncertainty such as the coefficient of variation in the demand. How do you assemble all of the above information to develop a final performance-based design? Structural engineers use common sense when developing the final design of all structural members, including dynamic modification devices, which comprise the structural system to resist the wind and earthquake forces on the building. Estimation of the lateral force (wind/seismic) demands (expected values) and the capacity (strength) for each limit state are carried out. Then lateral force expected values are multiplied by the so-called load factor, used to develop final design lateral force. It is this load factor, with a value .1, that mathematically considers both the statistical variation in the lateral forces and the consequences of failure, in order to make a structural design to not fail, with an acceptable probability of failure, the level of safety, set by our building codes and standards. This same basic approach is used to calculate the design capacity, strength, of each limit state. In this case, we multiply the expected capacity of the structural member and system by a capacity reduction factor whose numerical value is ,1 and consider both the variability and the consequences of failure in our estimate of capacity.

Executive Summary

When dynamic modification devices are utilized, the expected demand from the lateral load is significantly reduced due to the energy absorption by the damper. Also because these devices are manufactured and tested on very exact quality standards, they have much less uncertainty in their capacity compared to, for example, concrete structural members in place. How do we communicate all this important information? All of these mathematics and computer run time are of no value unless structural engineers communicate with nonstructural engineers. Therefore, in structural reliability, a common communication language needs to be used between the structural engineer, the decision maker, and the projects’ structural engineering peer-review panel and building officials. The words just defined above, demand imposed on the structural system, capacity of the structural system to resist these demands, and limit states are some of the words we use in this communication. Why dynamic modification devices in tall buildings? In its most basic view, dynamic modification devices in the performance-based design of buildings are, as previously noted, like sponges that absorb the energy imported from the lateral load to a building’s structural system and for a specific amount of imported wind/ seismic energy result in the concrete, steel, or other materials having to absorb a smaller percentage of the input energy and therefore reduce structural damage or building motion. In addition and very important in design is that this energy absorption by dynamic modification devices is quantified in a much more predictable way that benefits the final confidence levels in the structural performance-based designs. The addition of manufactured energy dissipation devices to tall and super tall buildings is very beneficial because it reduces the forces, that is, the demands, on limit states, on structural members and structural system. Importantly, the addition of dynamic modification devices also enables to calculate these forces with a very superior level of confidence. In modern structural engineering, this reduction with confidence will always benefit clients either monetarily or emotionally. A very essential goal of a structural engineer is to bring architects’ vision into reality and make sure that dynamic modification devices can always be located in the building so that they meet the architects’ and decision makers’ visions and also absorb great amounts of energy without damaging the steel and concrete structural members.

5

CHAPTER

Introduction

2

CHAPTER OUTLINE 2.1 Aim of the Book ................................................................................................. 10 2.2 Terms and Definitions......................................................................................... 12 2.3 List of Symbols and Abbreviations....................................................................... 21 2.3.1 Abbreviations ...................................................................................21 2.3.2 Symbols ..........................................................................................21 2.4 History of Dynamic Modification Devices in Tall Buildings.................................... 28 2.5 Industrial Application of Damping Devices ........................................................... 30 2.5.1 Vibration Control ..............................................................................30 2.5.2 Vibration and Seismic Control of Industrial Machine Structures ............32 2.5.3 Vibration Isolation of Buildings and Rail Structures .............................34 2.5.4 Vibration Control of Large Industrial Chimney Stacks ...........................34

Tall buildings have become a solution to respond to the need for increased density in major cities around the world. The trend of the past few years has been to build taller and slimmer structures, as demonstrated by the tall buildings under construction or recently completed around the world. This has been possible, thanks to the latest advancements in high-strength materials and construction methods that have led to more efficient structural solutions (Ali and Moon, 2007). However, these lighter systems could lead to structures that are more prone to vibrations, which can cause discomfort, damage, and structural failure. Tall and slender buildings need to withstand a variety of external forces that are different from those of low-rise construction, and as a consequence, different structural solutions must be applied. Moreover, many major cities are threatened by a variety of extreme events such as earthquakes and strong winds. Therefore, design of tall buildings should take both static and dynamic loads into consideration. This could be accomplished by modifying stiffness and increasing damping (energy dissipation), while minimizing overall building weight. Building stiffness is related with the proper selection of a structural system. For example, tubes, diagrids, and core-supported outrigger structures are often considered preferable solutions to increase stiffness and to reduce displacement response, but they might pose problems for acceleration response. Alternatively, increasing building damping can be an option achieved by installing supplementary damping devices. Moreover, damping characteristics of the main structural system (i.e., inherent Damping Technologies for Tall Buildings. DOI: https://doi.org/10.1016/B978-0-12-815963-7.00002-6 Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved.

7

8

CHAPTER 2 Introduction

FIGURE 2.1 Possible supplementary damping systems for tall buildings.

damping) is largely uncertain until the building is complete (Smith et al., 2010). In contrast, damping provided by external devices can be estimated quite accurately, thanks to the extensive research conducted in the past decades, and the confidence in the performance of damping devices, making dampers a reliable solution in reducing vibrations in tall buildings. Damping systems can be subdivided into three main categories: (1) passive, (2) active, semiactive, and hybrid systems, and (3) seismic isolation (Fig. 2.1) (Kareem et al., 1999). Passive and seismic isolation systems have fixed properties, while active systems change their properties based on the load demand and require an external energy source to be activated. Therefore, even though they are more efficient, active systems are less common than passive systems due to economic and reliability constraints. Passive systems can be further divided into two subcategories: (1) materialbased dissipation systems (i.e., distributed dampers such as viscous and

Introduction

viscoelastic dampers) and (2) mass-based dissipation systems generating large differential movements from rigid motion (i.e., tuned mass dampers (TMDs) or turbulence (e.g., tuned liquid dampers (TLDs)) to convert motion into other energy forms (such as heat). The first category of passive dampers is typically an integral part of the primary structural systems that would be positioned in optimal locations (e.g., within bracing systems) to reduce the building’s dynamic motion. The second category of passive systems is frequently positioned at the top of a tall building, although it might be used in different elevations depending on the desired performance. Different from the second category of passive systems, which are tuned to work on a narrow range of dynamic behaviors (e.g., modes and periods), active systems perform more efficiently over a wider range (Connor, 2003). There are many different types of active devices, but the most prominent ones are active mass dampers (Kobori et al., 1991a,b) and active variable stiffness devices (Kobori and Minai, 1960). Contrary to passive dampers, active dampers are adjusted through a computerized control system. It is believed that further research on these devices might lead to an increase in their utilization. The utilization of active, semiactive, and hybrid systems relies on the extensive data from building research conducted by the several national foundations in the United States, Japan, and China. One major concern of their utilization in the past was reliability, but this has been overcome for most damping devices by adopting hybrid control systems based on the combination of passive and active control methods. As an example, the hybrid mass damper in the 300 m Yokohama Landmark Tower can be mentioned (Preumont and Seto, 2008). Moreover, with the introduction of the concept of resilient buildings, the utilization of this technology has been considered as an enhancement in building performance as it reduces damage and increases life safety. Seismic isolation systems are considered independent from the first two categories, since the main function of the system is to decouple the building response of the structure above the isolation level. There are two main device categories that belong to isolation systems: bearing and sliding. Their applicability to tall buildings has always been debatable, since to decouple the tall building motion from the ground would require really a flexible isolation system. However, in the last decade, several tall buildings have been equipped with isolation devices, and this is the reason why they will be reviewed in this book. Supplementary damping may also be used in existing tall buildings undergoing a retrofit. For example, the 54-story Shinjuku Center Building in Tokyo has been retrofitted with velocity-dependent oil dampers (which utilize flow resistance of oil with low viscosity) to overcome problems in the existing building’s structural capacity under long-period ground motions (Aono et al., 2011). When designing/implementing supplementary damping, some essential aspects to be considered are device maintenance and reliability. For this reason, health monitoring has become a relevant topic for the life-cycle behavior of structures with additional damping devices. Several papers have been published on this topic

9

10

CHAPTER 2 Introduction

(e.g., Tamura et al., 1995; Bashor et al., 2012). Moreover, it is important to consider the life-cycle cost analysis of buildings with supplementary damping in order to evaluate their cost-effectiveness (Chen et al., 2008; Hahm et al., 2013). For example, the benefit of supplementary damping devices in a high-rise steel moment frame building under earthquake demands was estimated by King et al. (2001), depending on specific thresholds of the additional costs coming from the installation of these devices (e.g., 10% added damping would be beneficial for a 30-year time period if the added cost is less than $4.9 million). See Section 3.6 for a detailed discussion on this topic.

2.1 AIM OF THE BOOK Given the importance that dynamic modification technologies have in the design, performance, and safety of tall buildings, the purpose of this book is to provide the reader with information about the most important tools to understand, study, and design tall buildings equipped with these devices. This book intends to reach an interdisciplinary audience, from building designers (e.g., architects, engineers, consultants, contractors) to professionals involved in the construction and maintenance industry (e.g., contractors, damping suppliers) and building occupants (e.g., building users, general public, public administration officials). Both building professionals and building users can benefit from this book because it provides background information as well as major technical details involved in damping technologies for tall buildings. This book reflects international best expertise and knowledge based on the current state of the art on the subject. The recommendations provided are not intended to supersede national codes or other design recommendations but rather to provide general information to be considered by anyone who deals with tall building dynamic modification technologies. The book is divided into 10 chapters that overview the main aspects of damping technologies utilized in tall buildings. Chapter 3: The importance and relevance of damping technologies for tall buildings is the main topic of this chapter. The basic principles of dynamic response of buildings are first presented. The possibilities for controlling building performance and improving building resilience are addressed. Also, the main advantage of utilizing damping technology, protecting the building and its users from extreme events, and providing occupant comfort are discussed. Moreover, sustainability and environmental benefits are addressed. Particular focus is given to the life-cycle assessment of buildings equipped with damping devices. Design alternatives to damping are listed and described for comparison purposes. Chapter target audience: building owners and users, architects, engineers, public administration officials, consultants, the general public, and contractors Chapter 4: This chapter reviews the major properties of damping devices and general modeling approaches. Their relative advantages and limitations are discussed for

2.1 Aim of the Book

both physical and behavioral characteristics. Major systems are reviewed and divided into three main categories: (1) passive, (2) seismic isolation, and (3) active, semiactive, and hybrid. In each category, potential applications of damping systems are identified, based on the type of excitation (e.g., wind and/or earthquake) and the possible building configurations or locations in which they may be incorporated (e.g., braces, top of the structure, foundation, etc.). Using a combination of different devices is discussed, as there could be a possible solution that utilizes the best features of multiple device types. For this reason, the chapter ends with a comparison of the different devices, with a summary of the major advantages and limitations of each. Chapter target audience: building owners, architects, engineers, public administration officials, and consultants Chapter 5: Recommendations for the dynamic modification device design are the challenge discussed in this chapter. First, current codes and standards available worldwide are reviewed, with a particular examination of US codes. Second, practical design aspects and design tools for structures equipped with dissipation devices are examined in order to understand the main parameters to consider in the design process. Each damping category (passive, isolation, active, semiactive, and hybrid) is reviewed, and its detailed step-by-step design procedure is presented. These are based on current building code requirements and additional recommendations found in available literature. With this information, the designer should be able to generate a preliminary design for a tall building with dynamic modification devices. In addition, a special section is devoted to existing buildings undergoing a retrofit with dynamic modification system as the major factor to enhance the performance of the building. The chapter ends with a discussion on dynamic modification device optimization strategies. It is important to state that step-by-step procedures provided are merely recommendations for preliminary design of buildings with added damping devices. The designer should always refer to the relevant codes for specific design requirements. Chapter target audience: engineers and consultants Chapter 6: Dynamic modification devices represent a quite novel approach to the design of tall buildings, even though their implementation in civil structures started more than 50 years ago. For this reason, most professionals involved in building design do not have a clear understanding of the correlations that these devices have on other building systems. This chapter provides an overview of the interactions between building systems and damping devices. In particular, architectural, vertical transportation, mechanical, and fac¸ade aspects of tall buildings are reviewed. Chapter target audience: building owners and users, architects, engineers, public administration officials, consultants, the general public, and contractors Chapter 7: This chapter discusses the major requirements for testing, inspection, and maintenance of dynamic modification devices. Testing represents the most important tool to validate design assumptions and to understand device reliability and performance. Code requirements are utilized as the basis of discussion,

11

12

CHAPTER 2 Introduction

together with the supplier’s warranty. General specifications for the transportation, construction, and care of devices are also explained due to requirements that may be different from standard construction techniques. The chapter ends with a review of major building rating systems (i.e., resilience-based earthquake design initiative and US Resiliency Council) and how performance could be enhanced when dynamic modification systems are utilized. Chapter target audience: building owners and users, architects, engineers, public administration officials, consultants, the general public, and contractors. Chapter 8: Case studies of tall buildings equipped with dynamic modification devices provide good examples on the myriad of benefits and consequences of adding damping technology to building structures. Several case studies explain not only technical details about the design but also the processes and special features involved in the development of buildings equipped with dynamic modification devices. Examples are given for designed, built, and retrofit projects. Chapter target audience: building owners and users, architects, engineers, public administration officials, consultants, the general public, and contractors Chapter 9: Previous chapters reviewed the current state-of-the-art dynamic modification technologies in tall buildings, but this chapter introduces what could be the future of damped systems. This is particularly significant for understanding potential research areas for advancement of the technology. Chapter target audience: building owners and users, architects, engineers, public administration officials, consultants, the general public, and contractors Chapter 10: This chapter summarizes the data trends found for tall buildings with dynamic modification devices and provides final remarks on the topics reviewed in the whole book. Chapter target audience: building owners and users, architects, engineers, public administration officials, consultants, the general public, and contractors

2.2 TERMS AND DEFINITIONS Action: Internal moment, shear, torque, axial force, deformation, displacement, or rotation result of a degree-of-freedom displacement: designed as force or deformation controlled Deformation-controlled action: An action on a member or element permitted to exceed the yield value of the element Force-controlled action: An action on a member or element not permitted to exceed the nominal strength of the element Active fault: Fault being considered active by the authority having jurisdiction Acceptance criteria: Acceptability limits for different behaviors (e.g., strength, drift, inelastic demand) for a given performance level. Aspect ratio: See slenderness ratio. Backstay effect: Transfer of shear forces from the lateral system of the tower to additional elements that exist within the podium, through one or more floor diaphragms.

2.2 Terms and Definitions

Base level: Level of the structure at which lateral forces are transferred to the ground; for structures without a basement, it corresponds to the foundation level; for base-isolated structures, it corresponds to the midheight isolation. Base shear: Total lateral force or shear just above the building’s base level induced by lateral forces (wind or earthquake). Beam: Member mainly subjected to flexural loads. Building: Any structure providing support or shelter intended for occupancy use Building drift ratio: Building displacement at roof level divided by the height of the building above the structural base. Capacity design: For structures subjected to earthquake forces, the required element strength is defined by the forces corresponding to the expected capacity of the confined regions where elements yield. Characteristic earthquake: For an active fault, the estimate of the maximum earthquake magnitude that could occur, in any case, not less than the maximum that has historically occurred. Column: Member subjected mainly to compressive axial loads. Coupling beam: Beam that combines structural walls or any other lateral loadresisting members acting in-plane. Damping: The phenomenon that results in a decrease in the amplitude of oscillation through energy dissipation in a dynamic system. Additional/added/supplementary damping: Damping added to the structure by devices provided for that purpose Aeroelastic or aerodynamic damping: Damping related to the structure moving in a fluid (air in this specific case) Coulomb or friction damping: Related to energy dissipated by friction between two surfaces Damping constant: For viscous damping, coefficient used to represent damping magnitude in a system or device, expressed as a function of velocity Critical damping: Smallest value of damping that allows the system to return to steady-state position without oscillating Effective damping: Value that determines how much energy loss occurs in the system under analysis Equivalent viscous damping: Approximation of energy dissipation, coming from any source, in terms of viscous damping; this simplification is commonly used in structural engineering for simplicity of modeling damping in structural analysis Hysteretic damping: Energy dissipated due to the inelastic behavior of structural members Modal damping: Damping associated with a particular mode of vibration of the structure Inherent/intrinsic damping: Building damping related to building internal properties such as material, structural joints, soil
structure interaction, and nonstructural elements

13

14

CHAPTER 2 Introduction

Radiation damping: In soil
structure interaction, it is defined as the energy dissipated through radiation of waves away from the foundation Damping ratio: The ratio of system damping divided by critical damping Rayleigh damping: Damping expressed as a proportion of the mass and stiffness of the system Total damping: Combination of inherent, hysteretic, aerodynamic, and additional damping Viscous damping: Velocity-dependent energy losses as occurring in liquid lubrication between moving parts or in a fluid pushed (e.g., with a piston) through an orifice Damping system: All the additional or supplementary devices and related members, linkage, and connections added to a structural system to increase energy dissipation in the building. Distributed: Damping system in which the supplementary devices are spatially distributed throughout the entire building system (whether viscous, viscoelastic, metallic, etc.) Discrete: Damping system in which the supplementary devices are located in only a few locations in the building system (whether tuned mass dampers, tuned liquid dampers, etc.) Damper: Device or detail that converts kinetic energy (motion) to thermal energy (heat). Fluid viscous damper: Energy dissipation from a piston pushing a fluid through an orifice Friction damper: Energy dissipation through friction between surfaces Lead damper: Energy dissipation from pushing lead through an orifice Mass damper: Energy dissipation from behavior driven by differential motion between the building and additional mass moving out of phase (lagging behind) Metallic-yielding damper: Energy dissipation provided by yielding of the damper material Oil damper: Energy dissipation from a piston pushing oil with low viscosity through an orifice Shape memory alloy damper: Energy dissipation from inelastic behavior of special memory alloys Steel damper: A device with steel material that dissipates energy through its plastic deformation Tuned mass damper (TMD): A mass, spring, and damper system usually placed at the top of the building Tuned liquid damper (TLD): Transfers energy from the building to the liquid. Usually, it is a container placed at the top of the building. In case, it is a vertical U-shaped tank, the system is a tuned liquid column damper (TLCD) Viscoelastic (VE) damper: Device in which the cyclic deformation and subsequent relaxation of viscoelastic material produces energy dissipation

2.2 Terms and Definitions

Deformation Elastic deformation: Member deformation that is proportional to load and returns to zero when load is removed Inelastic deformation: Member resistance that reaches a yield limit when it is deformed, and deformation does not return to zero when load is removed Design basis earthquake (DBE): Earthquake level defined by the code response spectrum; it can be estimated as two-thirds of the maximum considered earthquake (MCER) excitation. Design engineer: Qualified person who designs the structural elements under the design loads. Design strength: Required strength times the load factor. Diaphragm: All the horizontal structural members (roof, floor, bracing, etc.) that are needed to transfer lateral forces to the vertically oriented lateral load-resisting system Displacement (for a supplementary damping device) Design displacement: The design earthquake lateral displacement, excluding additional displacement due to actual and accidental torsion; in the case when these are included, it is called total design displacement maximum displacement: The maximum earthquake lateral displacement that the dissipation system needs to be designed for, excluding additional displacement due to actual and accidental torsion; in the case when these are included, it is called total maximum displacement Displacement-dependent damping device: A device that dissipates energy as a function of the displacement between device ends (i.e., linear spring, like a rubber base isolator). Displacement restraint system: A group of elements that restrain the lateral movement of the isolated structure. Dissipation: See energy dissipation. Ductility: The capacity of a cross section, a structural member, or an entire structure to sustain loads and dissipate energy, through inelastic deformation under the design seismic event, without strength loss or failure. Dynamic test: A test in which actual working conditions of a system’s dynamic properties (e.g., natural frequency, damping ratio, and dynamic response) are measured. Dynamic modification systems: All the devices that change the dynamic properties of a building (damping devices, seismic isolation, etc.). Energy dissipation: Transformation from mechanical energy (motion) of the system into heat, which is dissipated to the surroundings by convection, conduction, or radiation. Energy dissipation device: A nonload-bearing device specifically designed to dissipate energy. Energy dissipation system: A collection of energy dissipation devices, related to supporting framing and connections.

15

16

CHAPTER 2 Introduction

Effective stiffness: Force divided by displacement for the system at the design value. Essential facilities: Structures that need to remain operational after a major event (e.g., wind, snow, and earthquake). Factory test: A testing program for mass-type dampers, held prior to installation, to approve that device properties comply with the design parameters; dynamic vibration tests and techniques are included. Fatigue endurance limit: The lowest stress at which the component will fail. For any values below the limit, the component will not fail regardless of the number of cycles it undergoes. Flat slab: A two-way, column-supported continuous slab without beams. Forced vibration test: A test conducted to identify a structure’s dynamic properties (frequency and damping ratio) based on measuring the response under a known excitation force, for example, harmonic (harmonic vibration test). Frame: A structural system that resists loads with a combination of beams, columns, and braces. Braced frame: A vertical truss that resists lateral loads, for example, concentrically braced frame or eccentrically braced frame Moment frame: Framing systems that provide stability and resistance to lateral loads mainly through shear and flexure Dual system: A system that resists lateral loads using a combination of moment frames and structural walls or braced frames Free vibration test: A test conducted to identify natural properties (frequency and damping ratio) of a system based on measuring the natural response of a structure to some impact or displacement. Fundamental period: A building’s natural period for the direction of interest that has the lowest frequency or longest period. Gravity system: A structural system intended to resist gravity loads; it may have limited ductility, and it is assumed to provide little or no lateral load resistance. Human-powered vibration test: A test conducted to identify a structure’s dynamic properties (frequency and damping ratio) based on measuring the response under a human-powered excitation force. Hysteresis: The relationship between material stress and strain, or member force and deformation, during cyclic loading; it represents the behavior history of a physical system. Importance factor: Numerical value for adjusting demands or acceptance criteria to reflect different reliability requirements based on building function or occupancy type. Interstory drift: The difference in horizontal deflection between two adjacent floor levels that are one story apart. Isolation interface: A plane- or surface-defining transition location between the isolated and the nonisolated structural parts of a building.

2.2 Terms and Definitions

Isolation system: All the structural elements (including isolation devices, wind restrain system, and dampers) that accommodate the isolation forces and displacements. Isolation device: A load-bearing element with small horizontal and large vertical stiffness capable of accepting large deformations under the design in seismic ground motions. Life-cycle assessment: The procedure to evaluate environmental impact of a product from fabrication and construction to demolition and disposal. Limit state: The state beyond which the design criteria are not met or satisfied. Serviceability: The state above which service requirements for a structure or structure member are no longer met Ultimate: The state that relates to structural failure and corresponds to the maximum load-carrying capacity of the system Load and forces Dead load: The effect of mass of the permanent component (structure, partition, finishes, etc.) under the influence of gravity Seismic load: The force used when designing to resist earthquake effects Load factor: The factor that takes into account the deviation from actual and nominal load due to uncertainties Load path: The load on a continuous structural system from the point it is applied or generated to the supporting system where it is resisted (e.g., soil) Lateral load: Load mainly acting in a horizontal direction, for example, wind and earthquake Live load: Load resulting from the occupancy or use of a structure, typically due to gravity Nominal load: Load magnitude specified by codes and/or standards Wind load: The force assumed to simulate the effect of wind acting on a building Multiple degrees-of-freedom system (MDOF): A system whose motion can be described only by multiple coordinates that are functions of time. The system can be modeled as linked, multiple mass
spring
damper subassemblies. Nonlinear static procedure: An analysis procedure of the structural system that takes into consideration the nonlinear properties of structural members and utilizes an incremental loading input on the building; sometimes termed a “pushover” analysis. Nominal strength: Structural capacity based on material-specified properties Maximum considered earthquake (MCER): The most severe or rare (a 2% probability of being exceeded in 50 years) earthquake that could impact the structure based on the site conditions. Member: Primary load-bearing component of the building structure, such as wall, beam, column, slab or connection, and damping/isolation device Occupancy: The intended use of a structure.

17

18

CHAPTER 2 Introduction

Online hybrid test: A test in which the whole system is modeled experimentally and is controlled by a computer with direct numerical time integration to solve equations of motion. If a part of the system is considered experimentally, the term substructures online (hybrid) test is used. P-delta effect: Structural actions resulting from axial loads (P) laterally displaced (by delta, δ) from their initial alignment (e.g., P as gravity loading and delta as building or story sidesway). Panel deformation: In-plane shear racking deformation of a wall panel, defined as the difference between the interstory drift and the local floor slope Peak ground acceleration: The largest ground (absolute) acceleration recorded by an accelerogram at a site during a particular seismic event. Performance level: Features like “immediate occupancy,” “life safety,” or “collapse prevention” that are used to establish a set of specific targets for accepting member forces and deformations; performance requirements shall be discussed between the client and engineer. Performance-based design: The process in which several performance levels need to be satisfied based on different demand levels; the choice of the performance levels is based on the discussion between the client and engineer. Plastic hinge region: The zone of the member in which significant inelastic deformation can develop under flexural actions. Posttensioning: Prestressing techniques in which the tendons are tensioned after the concrete has hardened. Prestressed concrete: Concrete in which internal compressive and flexural stresses are induced by pretensioning or posttension in order to increase the load capacity. Preinstallation test: A test that may include any kind of testing approaches in which the performance of a device is verified prior to the final installation within buildings. Pretensioning: Prestressing techniques in which the tendons are tensioned before the concrete has been placed, typically for precast, prestressed concrete elements. Production test (or factory production control test (European standard), or performance test (Japanese and Chinese standards)): Test performed on products/devices prior to installation in buildings to demonstrate that the nominal design parameters and damping/isolation system characteristics (force
velocity
displacement behavior) are between limit values (minimum and maximum) determined during the design process. Prototype test (or type test (European standard) or material test (Japanese standard)): Test conducted prior to product/device production for several purposes: estimating damping/isolation system characteristics (force
velocity
displacement behavior) and establishing the robustness of individual devices under predominate excitation, for example, earthquake. The nominal design properties determined are based on the data obtained from these tests.

2.2 Terms and Definitions

Qualification tests: Tests on isolation devices conducted by the manufacturer to be submitted to the registered design professional to quantify the variations in properties due to various effects such as heating, temperature, aging, etc. Quality control: The testing and inspection program in which quality and consistency of the manufacturing process is validated. Measurements include the as-built properties of the damping/isolation systems, prior to installation. The actions to take in this program are to be agreed between the designer and manufacturer. Quasi static test: A test device in which the loading/displacement rate is not significant, that is, the rate is slow. Racking deformation: Average panel deformation for a story. Redundancy: Property of a structural load-resisting system that has an alternative load path for transferring forces in case of local element failure. Required strength: The largest force(s) for which an element must be designed Resilience: A measure of how quickly a system recovers from shock. Resistance factor: Factor with a value less than 1.0 to reflect the probability of actual capacity varying from nominal capacity. Response
history procedure: Analysis to determine building responses (e.g., forces and deformations) over time to a forcing function that varies over time. Linear: Response
history procedure in which the structural system behavior is linear (elastic) Nonlinear: Response
history procedure in which the nonlinear behavior of the structural system is explicitly modeled; it also includes the case in which geometrical nonlinearities are present Response spectrum: Plot of the maximum or steady-state response quantities (i.e., displacement, velocity, or acceleration) of a series of equivalent single degree-of-freedom (SDOF) oscillators of varying period or frequency and of a given damping, subjected to a particular force. In the case when the forced motion is an earthquake, it is called a seismic response spectrum. Spectral: Maximum response quantities (i.e., displacement (SD), velocity (SV), or absolute acceleration (SA)) determined for a series of equivalent SDOF oscillators Pseudo: Approximated values of the velocity (PSV) and absolute acceleration (PSA) spectral response quantities, determined from the spectral displacement (SD) and the circular frequency of the equivalent SDOF system Retrofit: Design and construction activities performed for improving the structural performance of an existing system. Return period (or mean recurrence interval): Inverse of the annual probability of events (either seismic or wind) that are equal to or greater than a specified value. Risk category: A building classification to determine the loads (e.g., flood, wind, snow, ice, and earthquake) based on the risk related to undesirable performance. Seismic force-resisting system: Those portions of the structural system designated to resist lateral forces from seismic events.

19

20

CHAPTER 2 Introduction

Seismic hazard level: Probability of occurrence of a given earthquake intensity in a given area and return period. Shaking table test: A testing procedure in which a special device (shake table) simulates earthquake inputs to determine the building or structure component response based on response of an appropriately scaled model. Shear wall: Solid vertical planar element used to resist the lateral loads in its plane as well as to provide structural stability. Shear-type model: Simplification of a structural model in which a mass is concentrated in each floor and connected to the adjacent ones through shear springs with simple hysteretic characteristics. Single degree-of-freedom (SDOF) system: A dynamic (time-dependent) element composed of a single mass
spring
damper assembly having only one degree of freedom. Site class: Soil classification of near-surface conditions at the site of interest based on standard requirements. Slenderness ratio: Building height divided by footprint width. Soil
structure interaction: Study of the ways that soil properties and behaviors affect the structural response of a building. Story: The structural portion between two floor surfaces at different elevations Drift: Difference in horizontal deflection between two levels Drift ratio: Drift divided by height between the levels considered Shear: Lateral force in the story as the summation of all the lateral forces coming from all the floors above Strength design: A method for structural design in which the member actions do not exceed the design strength. Structural system: A load-resisting system of the structure under consideration; the damping system is a part of the structural system. Substructure hybrid test: A testing approach for structures (e.g., building including damper/isolation) in which the inelastic part of the structure (damper/isolation) is analyzed experimentally and the elastic part (building system) is numerically analyzed using a computer; then, the overall structural response is determined by combining the responses of the two parts. Velocity-dependent damping device: A device whose force response is mainly a function of the velocity of differential movements between the two ends. This is also called “viscous” damping. Wall: A vertical planar structural element with thickness much smaller than plan length or height. Wind restraint system: All the structural elements provided to resist lateral movement of the isolator units under wind load. Wind tunnel test: A procedure using scale models loaded by fan-driven airflow to predict responses of a structure, structural components, and cladding to actual wind storm conditions. Yield stress: Stress limit at which the stress
strain proportionality is deviating, the resisting force does not increase as much for further deformations, and upon release of load, residual deformations remain.

2.3 List of Symbols and Abbreviations

2.3 LIST OF SYMBOLS AND ABBREVIATIONS 2.3.1 ABBREVIATIONS ADAS CQC FNA HMD LED MDOF MR NLTHA PO PBD RC RMS SDOF SMA SHM SSI SYC TLCD TLD TMD

added damping
added stiffness systems complete quadratic combination fast nonlinear analysis hybrid mass damper lead extrusion damper multiple degrees-of-freedom system magnetorheological damper nonlinear time history analysis performance objective performance-based design reinforced concrete root mean square single degree-of-freedom system shape memory alloys structural health monitoring soil
structure interaction scorpion-yielding connector tuned liquid column damper tuned liquid damper tuned mass damper

2.3.2 SYMBOLS 2.3.2.1 Subscript i j m N Nd N er Nm aM , aK aθ aexc a exc aθ Ag As Ash ATLCD At Aw Au Aled Bζ

degree of freedom or story level damper mode number of degrees of freedom number of dampers number of earthquake records number of mode shapes proportional constants for mass and stiffness angular acceleration normalized excitation amplitude normalized steady-state amplitude angular acceleration gross area of cross section area of nonprestressed tension reinforcement total shear area cross-sectional area of tuned liquid column damper torsion factor cross-sectional web nominal area cross-sectional area of U-damper annular area around the shaft in lead extrusion damper damping reduction factor

21

22

CHAPTER 2 Introduction

bd bd0 bdk bdTL Cda c cb cd cd;eq ceq cv cv;i cve cs;eq ccr cij cd cd;eq cN C  C Ci C Cm CR CS C Ck DI DL Dmax;m Dy D1 dc , dt DRc , DRt d ðxÞ d0 dled dc dk dTL e E

width or strip width in steel slit damper or metallic damper plate constant width in metallic damper (or width of yielding finger at its base) width of knee bracing system width of tuned liquid damper tank deflection amplification factor damping constant base slab damping damper damping constant equivalent damping coefficient equivalent viscous damping coefficient in structure viscous damper constant damping constant at the ith story viscoelastic damper damping equivalent damping constant in friction damper critical damping force associated with the ith degree of freedom due to unit velocity along the jth degree of freedom (damping influence coefficient) tuned mass damper damping coefficient tuned liquid damper/tuned liquid column damper equivalent damping coefficient nonlinear damping coefficient in tuned liquid column damper damping matrix generalized damping matrix of base-isolated structure diagonal matrix of damping coefficients diagonal damping matrix generalized damping corresponding to the mth mode risk coefficient seismic response coefficient constant to estimate higher mode damping ratio (see Eq. 3.40) stiffness coefficient calibrated from experiments for steel slit dampers damage index dead load maximum displacement corresponding to the mth mode yield displacement fundamental mode roof displacement DmD : roof displacement for the mth mode interstory drift corresponding to DRc and DRt , respectively. ratio of the damage cost to the replacement value for damage threshold at 50% and 0.5%, respectively. plate variable depth in metallic damper plate constant depth in metallic damper effective diameter in LED damper diameter in circular plate damper depth of knee bracing system depth of TLD liquid eccentricity modulus of elasticity

2.3 List of Symbols and Abbreviations

EC ED EH EL ES0 Es EN f fm f1 fd f d;opt fd fy fyopt f0 fmax f~ F ðtÞ Fa ðtÞ Fv ðtÞ F0 FI ðtÞ FL Fd ðtÞ FS ðtÞ Fy ðtÞ FIi ðtÞ Fdi ðtÞ FSi ðtÞ Fi ðtÞ Fm ðtÞ FðtÞ F d;max F d;min Fy Fdv Fd;ve Fd;led Fds FN FTP F g

concrete modulus of elasticity dissipated energy per cycle hysteretic energy earthquake load strain energy steel modulus of elasticity compressibility of silicone fluid natural frequency natural frequency corresponding to the mth mode natural frequency corresponding to the first mode frequency of isolated dampers optimal frequency ratio between TMD and main structure frequency ratio between isolated damper and main structure element yield strength optimum activation load characteristic strength total force of structure force ratio between damper and structure stiffness applied dynamic force (due to wind or earthquake) short period (at 0.2 s time periods) site coefficient for adjusting spectral response acceleration long-period (at 1.0 s time periods) site coefficient for adjusting spectral response acceleration amplitude of applied load inertial force flood load for load combination damping force spring force yield force inertial force associated with the structure’s ith degree of freedom damping force associated with the damper’s at ith degree of freedom spring force associated with the structure’s ith degree of freedom applied dynamic force associated with the structure’s ith degree of freedom (due to wind or earthquake) modal dynamic force associated with the structure’s mth mode of vibration force vector maximum force of the hysteresis loop minimum force of the hysteresis loop yield force damping force of viscous damper damping force of viscoelastic damper extrusion force of LED damper damping (slip) force of friction damper applied normal force tension force in pendulum TMD control force vector in active-based dampers ground acceleration

23

24

CHAPTER 2 Introduction

gc , gl , gQ , gR , and gV Gah Gav GE GC Geq Gef G1 , G2 , G3

H h hk hkf hkw hsyc hc hs hi HL HTL Ie Ig Is Iz J1 J2 J k kij

k0 kd k0p keff keq;p kve kE k~ K  K K Km Kdir

gust effect factor parameters for ASCE 7-10 (ASCE, 2010) wind load procedure horizontal geometric amplification factor vertical geometric amplification factor shear storage modulus of the viscoelastic material shear viscous damping constant equivalent shear modulus gust effect factor time-dependent control gain factors in active-based dampers for displacement, velocity, and excitation, respectively (Soong and Spencer, 2000) building height thickness thickness of knee bracing system flange thickness of knee bracing system web thickness of knee bracing system thickness of yielding finger in scorpion-yielding connector damper circular plate damper thickness shear thickness height above the isolation level hw : web thickness in steel slit damper heat load height of tuned load damper tank importance factor cross-sectional gross moment of inertia steel section moment of inertia intensity turbulence location matrix of the action of control force vector in active-based dampers location matrix of the action of external load vector in active-based dampers ASCE 41-13 (ASCE, 2013) force reduction factor stiffness force associated with the ith degree of freedom due to unit displacement along the jth degree of freedom (i.e., stiffness influence coefficient) initial stiffness damper or isolation stiffness postyield stiffness effective stiffness pendulum-tuned mass damper stiffness viscoelastic damper stiffness elastic stiffness stiffness ratio between damper and structure stiffness stiffness matrix generalized stiffness matrix of base-isolated structure diagonal stiffness matrix generalized stiffness corresponding to the mth mode wind directionality factor

2.3 List of Symbols and Abbreviations

Kzt LTM LTL LMR L L1 L2 Lδ Lled Lsyc l0 lk m mb md mi mij M  M M Mm Mp Mpo Nj N Ns Nm N Nd ns Ops p P qn ðtÞ q QE QCE QCL ra ru rfp rα R Rnew Rold RG rTL

topographic factor length of pendulum-tuned mass damper length of tuned liquid damper/tuned column liquid damper tank length of magnetorheological damper building width LL : live load building dimension along the widest vibration direction building dimension along the narrowest vibration direction factor accounting for logarithmic decrement of structural damping length of the sliding part of the shaft in lead extrusion damper length of yielding finger in scorpion-yielding connection damper strip length in steel slit damper length of knee bracing system mass base slab mass isolated damper mass concentrated mass of level (floor) i force associated with the ith degree of freedom due to unit acceleration along the jth degree of freedom (i.e., mass influence coefficient) mass matrix generalized mass matrix of base-isolated structure diagonal mass matrix generalized mass corresponding to the mth mode plastic bending moment plastic bending moment at midheight of metallic damper number of identical viscous dampers with the same damping cv;j number of degrees of freedom in a multiple degrees of freedom system number of building stories number of mode shapes blow count number number of dampers for the ith story or number of tuned mass dampers number of strips in steel slit damper amplification factors of pseudostatic response wind pressure axial load time-dependent displacement corresponding to the nth mode velocity pressure earthquake load for basic load combination expected material strength lower bound of material strength design acceleration reduction level (see Eq. 3.62) radius in U-damper radius friction pendulum isolator postyielding ratio rt : distance from the center of torsion response modification factor new response of structure old response of structure resonant factor for gust effect tuned liquid damper depth ratio

25

26

CHAPTER 2 Introduction

su SA SD SDS SD1 SL SS S1 SDm SAm t T Tb Tn T~ T0 TL TP u(t) u_ ðtÞ u¨ðtÞ u¨g u¨g0 ug ub udy uh ur ur;opt u0 ud ud0 u_ d ud;max ud;min ust U ðt Þ U_ ðtÞ U€ ðtÞ U t ðt Þ  U Vb Va V Vp v vs W

undrained shear strength spectral acceleration spectral displacement design spectral displacement at 0.2 s time period design spectral displacement at 1.0 s time period snow load for load combination spectral acceleration at 0.2 s time period spectral acceleration at 1.0 s time period spectral displacement corresponding to the mth mode spectral acceleration corresponding to the mth mode time natural period base slab period natural period corresponding to the nth mode temperature period as 0.2 SDS /SD1 long-period transition parameter pendulum TMD period dynamic translational displacement dynamic translational velocity dynamic translational acceleration ground acceleration ground acceleration amplitude ground displacement base slab displacement yield displacement of damper hysteretic component displacement function relative displacement between damper and structure optimal relative displacement between damper and structure amplitude of harmonic motion damper displacement damper displacement amplitude damper displacement rate (velocity) maximum displacement of the hysteretic loop minimum displacement of the hysteretic loop static lateral displacement dynamic displacement vector dynamic velocity vector dynamic acceleration vector total displacement vector generalized displacement vector of base-isolated structure base shear base shear above the isolation level shear force plastic shear force vertical displacement shear wave velocity weight

2.3 List of Symbols and Abbreviations

WL WP x y z

wind load for load combination weight of pendulum-tuned mass damper tip displacement amplitude of vibration eccentricity from center of rigidity to the element equivalent height for wind pressure definition

2.3.2.1.1 Greek symbols α α1 , α2 α, γ αp0 β χ δmax δu Δ E1 ,E2 φm φim φrj;m Φ φ γ1 , γ2 , γ3 γ γS γE γC γ_ S ΓðÞ Γm η ηs ηi ηd κ λ Λhd μ μ μD μM μled μs μfp

viscous damper velocity exponent parameters depending on the number of stories and primary construction material to calculate intrinsic damping (see Eq. 3.48) parameters related to building characteristics (see Eq. 3.49) postyield stiffness ratio in bilinear hysteretic loop model constant parameter (see Eq. 3.59) constant in lead extrusion damper maximum deformation ultimate deformation of the element story drift independent response parameters of self-centering systems mode shape corresponding to the mth mode ith modal element of the mth mode vector relative displacement between two ends of the jth damper corresponding to the mth mode modal matrix normalized deflection mode shape at the level of the absorber parameters depending on primary construction material shear strain in viscoelastic damper shear strain of viscoelastic material elastic shear strain viscous shear strain shear strain rate gamma function modal participation factor for the mth mode loss factor of viscoelastic material screen loss coefficient of tuned liquid damper damper proportionality factor along the building height tuned column liquid damper head loss coefficient knowledge factor property modification factor hysterically damped system’s parameter mass ratio ductility design earthquake ductility maximum credible earthquake ductility friction coefficient between the lead and steel shafts friction coefficient of friction dampers friction coefficient for friction pendulum

27

28

CHAPTER 2 Introduction

θ θk ρ ρF σ σg τS τE τC ω ω ωb ωb ωe ωd ωD ωm ωDm ωr ωn Ω2 Ω0 ψ Ψ ΨTL ζ ζ eff ζ eff ;opt ζ eq ζ eq;m ζ cr ζi ζd ζa ζh ζT ζm ζ im ζ Tm ζ d;opt

angle between brace and horizontal (vertical) direction angle between diagonal and knee elements density fluid density ratio between excitation frequency and system frequency ratio between ground frequency and system frequency shear stress of viscoelastic material elastic shear stress viscous shear stress circular (angular) frequency circular frequency of Kelvin element active system-effective bandwidth base slab frequency excitation frequency tuned mass damper frequency damped frequency frequency corresponding to the mth mode damped frequency corresponding to the mth mode frequency corresponding to the rth mode structure’s natural frequency spectral matrix of eigenvalue problem overstrength factor phase angles between response and excitation efficiency tuned liquid column damper efficiency damping ratio effective damping ratio optimal effective damping ratio equivalent damping ratio equivalent damping ratio corresponding to the mth mode of vibration critical damping ratio intrinsic damping supplemental damping aerodynamic damping hysteretic damping total damping damping ratio corresponding to the mth mode intrinsic damping corresponding to the mth mode fraction of total damping for the mth mode optimal supplemental damping ratio

2.4 HISTORY OF DYNAMIC MODIFICATION DEVICES IN TALL BUILDINGS Damping devices were first introduced in war applications in the late 1800s to suppress cannon and gun recoils. Later, damping was introduced in the

2.4 History of Dynamic Modification Devices in Tall Buildings

automotive industry (in the 1920s and 1930s) through the use of shock absorbers allowing drivers to have more comfortable rides. Subsequently, it was introduced in the general mechanical industry. Experimentation in the building industry started much later—during the 1950s. However, the first attempt at damping in buildings was made in 1870, with the patent for base isolation filed by Jules Touaillon for a building built on steel balls that rolled inside shallow dishes. In 1876, John Milne at the University of Tokyo also did some experiments for building on piles with steel balls. In 1909, J.A. Calantarients proposed a lubricated free joint that a building could sit atop. Despite these early experiments, research examining dynamic modification devices (passive, isolation and active, semiactive, and hybrid) did not begin in earnest until after 1950. Often, building research is funded and initiated as a response to the effects of a major natural disaster. Indeed, some of the major problems with conventional construction and the possible benefits of adding damping were brought up immediately following the earthquakes in Loma Prieta, California (1989); Northridge, California (1994); and Kobe, Japan (1995). One of the first major studies on the utilization of passive energy dissipation devices was carried out in New Zealand by Kelly et al. (1972) and Robinson and Greenback (1976). The major development push was carried out in the late 1980s and early 1990s (Kelly, 2001a, b) with several research efforts carried out by many companies (Hanson et al., 1993). Friction devices started to be developed by Sumitomo (Aiken and Kelly, 1990) and Pall Dynamics (Pall et al., 1987). Metallic system development started from triangular or hourglass shapes (Kelly et al., 1972; Tyler, 1978; Skinner et al., 1980; Stiemer et al., 1981). The application of viscoelastic materials to buildings started in 1969 with the work by Mahmoodi (1969), followed by the work of 3M company that provided the double-layer shear damper used in the dampers of the World Trade Center in New York City (Mahmoodi et al., 1987). More work was done in this direction, carried out on full-size steel frames (Lin et al., 1988; Aiken and Kelly, 1990). Fluid viscous dampers have been utilized extensively in the military and aerospace industry, and the transfer of this knowledge occurred in the 1990s, after the Cold War. Taylor Devices was one of the first companies to try applying this technology to buildings (Constantinou et al., 1993). Despite the early work on base isolation, civil applications started to be developed in the late 1970s, with the development of multilayer elastomeric bearings (Naeim and Kelly, 1999). The first use of these bearings occurred in 1969 at an elementary school in Skopje, Macedonia (Staudacher et al., 1970). Following this, isolation devices were installed in many more buildings around the world (Naeim and Kelly, 1999). After the Kobe 1995 earthquake, the use of high-rise, base-isolated buildings became common in Japan. The majority of these buildings, to date, were higher than 60 m and shorter than 100 m, but there were some buildings up to 180 m high, mostly concrete residential condominium buildings (Becker et al., 2015). The applications of active structural controls have a more recent history. This comes from several coordinated research studies between Japan and the United

29

30

CHAPTER 2 Introduction

States, starting with the US Panel on Structural Control Research (US-NSF) in 1989 (Soong and Spencer, 2000). The major challenge was the integration of civil engineering with several other disciplines, such as computer science, data processing, control theory, material science, and so on. Some of the early work on control design for civil structures was carried by Suhardjo et al. (1990) and Inaudi et al. (1993). After the early application in the World Trade Center (Mahmoodi et al., 1987), the widespread application of energy dissipation devices started in the 1980s and included applications such as the John Hancock Tower, Boston (see Section 8.2.1.3), and the Citicorp Center, New York City (see Section 8.2.1.2), with a TMD. Currently, there is a wide range of damping systems available (and buildings in which they were implemented), as will be shown.

2.5 INDUSTRIAL APPLICATION OF DAMPING DEVICES Supplementary damping systems, in combination with vibration isolation devices consisting of helical coil springs or elastomer materials (i.e., rubber, polyurethane, etc.), reduce transmissibility from industrial and power generation machinery, and road and rail transit, and protect sensitive machinery such as MRI in hospitals and precision equipment in clean room operations.

2.5.1 VIBRATION CONTROL Vibration and seismic control of large machinery and structures is possible using an elastic element, such as a steel spring, or elastomeric bearing. However, depending on the degree of isolation required, the tuning ratio (i.e., the ratio between disturbing frequency and support frequency) is critical and should be maintained at a minimum tuning ratio of about 2. If there are short-duration pulses, impact is expected, or operation is close to resonance (tuning ratio of 1); high degrees of damping are required to reduce vibration transmissibility. For the design of a vibration control system, considering the machine and the foundation as one rigid mass is usually sufficient. This is valid even if the machine itself is elastic, as long as the stiffness of the supporting isolation system is much lower than that of the machine and its foundation. The applied dynamic force causing the system to vibrate can be harmonic (such as from a fan or generator), periodic (such as from a high-speed press), transient (such as from pile driving), or random (such as from an earthquake or a subway train). The force may come from inside the system or from external sources. An important aspect of the dynamic behavior of a system is the relationship of the inertial forces (the inertia of the system mass when accelerated) to the applied forces. If the inertia forces are in phase (acting in the same direction) with the applied forces,

2.5 Industrial Application of Damping Devices

the support forces are greater than the applied forces, and there is no vibration isolation; rather, amplification occurs. If the inertia and applied forces are out of phase (acting in opposite directions), then they counteract each other, and there is only a small reaction in the elastic element and dampers. Vibration isolation is more than an arbitrary selection of any elastic mount under a machine or structure; it is a careful design decision. Since natural frequencies (see Section 3.1) are key characteristics of a vibrating system and are defined by the spring properties, examining the system based on the correct natural frequencies (both translational and rotational) is crucial. For different types of excitation, such as shock excitation or periodic excitation, different design criteria may be applicable. Horizontal translation and rotation about the corresponding horizontal axis are uncoupled only if the spring support is on the same level as the center of gravity, as shown in Fig. 2.2A. Otherwise, these degrees of freedom are coupled (Fig. 2.2B). The resulting combined motion is called rocking about a lower or upper rocking center. The response of the system is a combination or superposition of a horizontal motion and a rotation about the corresponding horizontal axis through the center of gravity. Using an elastic element with the lowest possible natural support frequency will generally provide the greatest degree of vibration isolation, but it may result in larger system displacements from operational loads or external excitation (i.e., material handling, wind, etc.). Mode coupling (Fig. 2.2B) should also be checked as rotating equipment can easily excite rocking modes with relatively minor imbalance forces.

FIGURE 2.2 Mode shapes for spring-supported systems: (A) uncoupled and (B) coupled modes.

31

32

CHAPTER 2 Introduction

2.5.2 VIBRATION AND SEISMIC CONTROL OF INDUSTRIAL MACHINE STRUCTURES Rotating equipment and air-handling units in buildings are often bearing supported to avoid the transmission of operational vibrations through floors, walls, and mechanical, electrical, and plumbing systems in buildings. Fig. 2.3A is an example of spring-supported, air-conditioning systems. This principle has been used to support larger fans in industrial cement production and power plants (Fig. 2.3B). Adjusting the modal properties of the integrated system is also an essential part of seismic control strategies. Such strategies include: • • • •

modifying the shape of the fundamental mode increasing the fundamental frequency increasing the damping limiting the forces transmitted to the structure

In a typical turbine generator the condenser sits below the low-pressure turbine leading to a table-type foundation. Fig. 2.4A shows a possible layout strategy for a spring-supported turbine generator foundation, compared to a rigidly supported case. No structural connection between turbine generator foundation system and adjacent machine building is allowed. In the spring-supported case, shown in Fig. 2.4B, the substructure of the foundation is arranged on a continuous base plate. A vibration isolation of up to 99% of transmitted operating forces through the foundation systems allows a continuous base mat, eliminates the need for double-column construction, reduces logistic issues with independent foundation plans, and reduces piles by lowering localized soil pressure. It also reduces

FIGURE 2.3 Examples of (A) bearing supported fans in an air-conditioning system and (B) a power plant.

2.5 Industrial Application of Damping Devices

FIGURE 2.4 A typical cross section of turbine generator machine buildings (A) without spring supports and (B) with spring supports. Spring-supported machine substructures are fully integrated with the surrounding building.

FIGURE 2.5 Elastic support of the diesel engine.

the risk of groundwater ingress, reduces the relative displacement between the structures, and reduces the accelerations in the machine and surrounding structure by harnessing a mass damping effect (being provided by the isolators, either by high-damping rubber or other damping devices) during a seismic event. These principles are applicable for all types of machinery. Fig. 2.5 shows the elastic spring support of a large diesel generator, typical of a small power station, but which can also be found in hospitals or large office complexes. The same system provides vibration isolation as well as seismic protection of critical emergency power systems when implementing larger damping devices that can handle larger displacements and forces than those excited by operation vibration. Even substation equipment, such as transformers and capacitor banks, can be isolated for vibration and seismic control (Fig. 2.6).

33

34

CHAPTER 2 Introduction

FIGURE 2.6 Vibration isolation of (A) transformer with (B) isolators.

2.5.3 VIBRATION ISOLATION OF BUILDINGS AND RAIL STRUCTURES Concert venues, recording studios, hospitals, high-tech manufacturing facilities, and laboratories are all buildings affected by vibration. Depending on the frequency content, intensity, and the vibration type, isolation is beneficial in the structure or at the source of the vibration (i.e., machinery or rail transit). For example, when a rail is directly passing through elevated, interior areas of a building, the transmission of vibration can be drastically reduced by engaging the inertia of a floating slab track. Some examples of vibration isolation being used to reduce the transmission of vibration from rail transit into buildings are shown in Figs. 2.7 and 2.8. Vibration from street trams and subways can be effectively mitigated using the same principle of the floating slab track. Fig. 2.9 shows an example of a jackup floating slab track at various stages.

2.5.4 VIBRATION CONTROL OF LARGE INDUSTRIAL CHIMNEY STACKS The most common mode of failure in stacks is vibration due to vortex shedding. When vortex shedding occurs with a frequency in the range of the natural frequency of the stack, these low-damped structures (damping ratio 5 0.5% of critical damping or less) display a resonant response and significant frequent, cyclic displacements that increase the fatigue loading. To reduce the resulting load, the damping of the stack has to be increased, which can be achieved by several strategies (such as Scruton helix).

2.5 Industrial Application of Damping Devices

FIGURE 2.7 Floating bridges on helical springs in Xizhimen station, Beijing, China.

FIGURE 2.8 Floating trackbed with a 6-Hz spring support at the Cheonan station high-speed line, Seoul, Korea.

Damping elements such as pads or dashpot dampers can be installed at the bottom of a base plate and the foundation so that they act in compression. Additionally, damping tensioners can be installed on the top of the base plate, so when the bolt is in tension, the damping washer is compressed. This type of damping element system serves to increase the damping of the stack structure and suppress wind-induced vibration on the stack. The addition of a damping pad slightly increases the flexibility of the base, which can result in slightly higher lateral deflections when the stack is subjected to wind loadings. The additional damping that can be achieved is small as the relative displacements are also very small.

35

FIGURE 2.9 Cast-in-place isolator housings and casting of concrete and tram on raised floating slab track.

2.5 Industrial Application of Damping Devices

FIGURE 2.10 Ring-type TMD at a stack—cement factory in Arquata, Italy.

Another way to increase the damping is to use a TMD, which is added to the top portion of the stack, to suppress wind-induced vibration (or certain seismic vibrations) (Fig. 2.10). Such a device consists of a mass and springs, tuned to a natural frequency of the structure, to drive supplementary dampers. The TMD can take different forms, but the most common style is to have a cylinder that acts as a mass and is larger than the diameter of the stack. The mass is suspended on pendulum ropes or rods. Between the mass and stack, there are springs for the tuning and damping elements. The entire assembly is designed to a specific natural frequency of the structure so that it will provide damping to suppress vibrations. TMDs are the most effective solutions, but can also be costly. A cost-optimized solution is to use a TLD, that is, an annular box partially filled with a water
glycol mixture. The box is specially designed to maximize the damping effectiveness of the fluid held by the damper. This style of damper is unique in that it can also be used as a platform with the addition of handrails. The damper has no moving parts. The damping is provided by the internally occurring sloshing of the liquid. The additional damping that can be achieved with a TLD is smaller than that of a tuned mass system due to poor tuning accuracy and nonlinear behavior of the fluid.

37

CHAPTER

Damping considerations in tall buildings

3

CHAPTER OUTLINE 3.1 Basic Principles of Building Response ................................................................ 41 3.1.1 Single Degree-of-Freedom System......................................................41 3.1.2 Multiple Degrees-of-Freedom System .................................................47 3.1.3 Equivalent Viscous Damping..............................................................53 3.2 Damping in Tall Buildings................................................................................... 55 3.2.1 Damping Modeling ...........................................................................55 3.2.2 Intrinsic Damping.............................................................................57 3.3 Effect of Damping on Building ............................................................................. 73 3.3.1 Other Development Considerations.....................................................77 3.4 Tall Buildings Wind-Excited Motion..................................................................... 78 3.4.1 Building Wind Vibration ....................................................................79 3.4.2 Occupant Comfort ............................................................................80 3.4.3 Wind Deflection Criteria ....................................................................84 3.5 Tall Buildings Earthquake-Excited Motion............................................................ 85 3.5.1 Principles of Performance-Based Earthquake-Resistant Design for Tall Buildings ..............................................................................87 3.5.2 The Role of Damping in the Seismic Response Control of Tall Buildings .....89 3.5.3 Seismic Deflection Criteria ................................................................90 3.6 Environmental and Economic Considerations ....................................................... 90 3.6.1 Life-Cycle Assessment ......................................................................93 3.6.2 Cost Implication of Buildings with Dissipation Devices ......................101 3.7 Damping Technology Uncertainty and Robustness Performance ..........................102 3.8 Alternative to Damping Devices.........................................................................103

Motions of a building, due to wind and seismic loads, are traditionally controlled using mass and stiffness of the structure. However, for taller buildings, this becomes inefficient, and the introduction of supplementary damping can provide a more economical overall design. This chapter outlines a history of the use of supplementary damping within buildings and gives an overview of the issues to consider when contemplating it. Supplementary damping reduces the response of the structure by removing a small amount of energy from the structure during each cycle of sway. For example, in viscous dampers, this is accomplished by the application of a force Damping Technologies for Tall Buildings. DOI: https://doi.org/10.1016/B978-0-12-815963-7.00003-8 Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved.

39

40

CHAPTER 3 Damping considerations in tall buildings

between two components of the structure, which is opposed to the velocity of these components and deducts the input energy to the building. The greater the velocity and opposing forces are, the greater the supplementary damping is. The history of the first supplementary damping in high-rise towers remains relevant till today. During the design of the original World Trade Center towers in New York City (the United States), important questions were raised about the governing design criteria for occupant’s comfort levels in a windstorm (Mahmoodi et al., 1987). In 1964 no one had yet attempted to determine the limits of acceptable perceived motion, a dynamics issue that would be critical in such a tall building. The design team knew it was an issue they must address if the project was to be a success, but they did not have a basis on which the design could be judged (Mahmoodi, 1969). In conjunction with early wind tunnel testing for the towers, they conducted a double-blind testing program to better understand human perception of motion and to answer the following question: when would occupants perceive accelerations due to the building’s sway? A clinical trial was developed and performed in Oregon (Eskildsen, 1965a,b). Volunteer subjects were offered free eye exams and brought into what appeared to be an examination room with a chair, doctor, and several attendants. Once they were seated, the room, which was mounted on actuators, started moving with increasing accelerations until the patient complained. Results from this study indicated that the initial design of the towers would result in an uncomfortable experience for their occupants during the frequent windstorms which they would be exposed to. The conventional solutions of adding stiffness with additional steel framing and adding mass with additional concrete would be inefficient, so motion reduction solutions by supplementary damping were considered. Ultimately, the tower designs were modified to include over 10,000 viscoelastic dampers distributed throughout each tower, providing energy dissipation on a scale that had not been attempted before. Since this early application, multiple advances in wind tunnel modeling, damping behavior predictions, and damping strategies for buildings have been developed. Tuned mass dampers, sloshing dampers, outrigger damping systems, link beam damping systems, braced frame dampers, active control systems, and other systems have become an innovative strategy and industry-accepted solutions when stiffness and mass alone cannot economically meet the improved performance goal. A properly designed supplementary damping system can be a highly efficient way for a structure to meet its improved performance goals. For a very tall structure, supplementary damping may be a required component to meet some specific requirement. However, increasing damping is not a magic solution for all building dynamics issues, and it may not be applicable to all buildings. It requires an experienced, capable design team with a detailed understanding of the dynamic properties of each specific building and the additional damping system.

3.1 Basic Principles of Building Response

FIGURE 3.1 (A) An idealized SDOF system and (B) relative forces in equilibrium.

3.1 BASIC PRINCIPLES OF BUILDING RESPONSE The dynamic behavior of mechanical (structural) systems subjected to external excitations (wind and earthquake) is an important aspect in structural dynamics. Structural systems are usually idealized to two major categories depending on the complexity of structure: single degree-of-freedom (SDOF) systems or multiple degrees of freedom (MDOF) systems. Understanding the concept and theory of these systems is essential for the study of dynamic modification system in buildings. For this reason, basic principles of dynamics are briefly discussed next.

3.1.1 SINGLE DEGREE-OF-FREEDOM SYSTEM An SDOF system (Fig. 3.1A) is a mechanical system in which the motion is described using a single translational coordinate [e.g., uðtÞ] as a function of time. If the system is assumed to have linear viscous damping, then its response to an external (dynamic) force F(t) is completely defined by the mass (m), stiffness (k), and damping (c) of the SDOF system (see Fig. 3.1A). For the sake of simplicity, these physical properties are, respectively, defined as concentrated in single elements: rigid block (for m), spring (for k), and dashpot (for c).

3.1.1.1 Equations of motion Based on d’Alembert’s principle, equilibrium of the forces acting upon the mass of SDOF system can be described as (see Fig. 3.1B): FI ðtÞ 1 Fd ðtÞ 1 FS ðtÞ 5 F ðtÞ

(3.1)

where FI ðtÞ is the inertial force, Fd ðtÞ the damping force, FS ðtÞ the spring force, and F ðtÞ the applied dynamic force (due to wind or earthquake). As defined by d’Alembert’s principle, the inertial, damping and spring forces are expressed as follows, respectively, FI ðtÞ 5 mu¨ðtÞ

(3.2)

Fd ðtÞ 5 cu_ ðtÞ

(3.3)

41

42

CHAPTER 3 Damping considerations in tall buildings

FS ðtÞ 5 kuðtÞ

(3.4)

where u¨ðtÞ denotes the acceleration, u_ ðtÞ the velocity, and u(t) the displacement of the SDOF system mass, relative to the ground. Substituting Eqs. (3.2)(3.4) in Eq. (3.1), the equation of motion of the SDOF system can be formulated by mu¨ðtÞ 1 cu_ ðtÞ 1 kuðtÞ 5 F ðtÞ

(3.5)

It is well known that the sources of energy dissipation in a building do not necessarily behave in a linear viscous manner. However, it is common practice to assume linear viscous damping to be proportional to velocity as this kind of damping facilitates solution of the equation of motion (Eq. (3.5)) without the involvement of second-order differential equations of motion (Jacobsen, 1930). Moreover numerous investigations have shown that employing this model of energy dissipation allows to correctly reproduce the measured response of instrumented buildings responding elastically to seismic excitations (e.g., Beck and Jennings, 1980; Cruz and Miranda, 2017a; McVerry, 1980; Safak and Celebi, 1991).

3.1.1.2 Free vibration properties The damped SDOF system (Fig. 3.1A) is at a free vibration condition when f ðtÞ 5 0. Therefore dividing the equation of motion (Eq. (3.5)) by the system mass (m): u¨ðtÞ 1 2ζωu_ ðtÞ 1 ω2 uðtÞ 5 0

(3.6)

where ζ is the damping ratio (fraction of critical damping ccr 5 2mω) and ω is the circular (angular) vibration frequency (radian/second), respectively, expressed as follows: c c 5 ccr 2mω rffiffiffiffi k ω5 m

ζ5

(3.7) (3.8)

Note that, ω is conceptually different from the natural frequency f (cycles/second) of the system, which is defined by: f5

ω 2π

(3.9)

Moreover the natural period of the systems can be expressed by: T5

1 2π 5 f ω

(3.10)

Depending on the value of ζ with respect to unity, three different system categories can be defined: underdamped system (ζ , 1), critically damped system (ζ 5 1), and overdamped system (ζ . 1). Note that structural (building) systems

3.1 Basic Principles of Building Response

are almost always classified as underdamped systems (see Chopra (2012) for a detailed discussion).

3.1.1.3 Forced vibration properties If the applied force F ðtÞ in Eq. (3.5) is not null, the equation of motion represents a forced vibration condition. Different forcing functions (e.g., harmonic and arbitrary) conditions would lead to different solutions for the SDOF equation of motion. If the SDOF system is excited by a harmonic force, F ðtÞ 5 F0 sinωe t, where F0 and ωe are force amplitude and excitation frequency, respectively, the solution of Eq. (3.5) consists of a transient response (first term) and a steady-state response (second term) in the following equation (see Fig. 3.2) (see Chopra (2012) for further details):   uðtÞ 5 e2ζωt ðA1 cosωD t 1 A2 sinωD tÞ 1 ½A3 sinωe t 1 A4 cosωe t

(3.11)

where ωD is the damped frequency defined as: ωD 5 ω

qffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 ζ2

(3.12)

The constants A3 and A4 are given by A3 5

F0 h k

 2 1 2 ωe =ω  2 i2   2 12 ωe =ω 1 2ζ ωe =ω

FIGURE 3.2 Response of a viscously damped system to a harmonic force.

(3.13)

43

44

CHAPTER 3 Damping considerations in tall buildings

FIGURE 3.3 Steady-state response of a viscously damped system to a sinusoidal force with phase-lag angle of 90 degrees.

A3 5

  2 2ζ ωe =ω F0 h  i 2 2   2 k 12 ωe =ω 1 2ζ ωe =ω

(3.14)

The parameters A1 and A2 can be determined by substituting the given SDOF initial conditions [uð0Þ and u_ ð0Þ] in Eq. (3.11). Alternatively, the steady-state deformation in Eq. (3.11) can be rewritten in the following format: uðtÞ 5 u0 sinðωe t 2 φÞ

(3.15)

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   where u0 5 A23 1 A24 is the response amplitude and φ 5 tan21 2 A4 =A3 is the phase angle that describes the extent that the response leads the applied force (Fig. 3.3). Subsequently, the deformation response factor can be found as: Rd 5

u0 1 5 rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi h  ðust Þ0 2 i2   2 1 2ζ ωe =ω 12 ωe =ω

(3.16)

where ðust Þ0 5 F0 =k is the static deformation under a constant force F0 . Moreover the phase angle can be expressed as: 21

φ 5 tan

  2ζ ωe =ω  2 1 2 ωe =ω

(3.17)

Similarly, the displacement dynamic response factor of Eq. (3.16) and the velocity (Rv ) and acceleration (Ra ) dynamic response factors can be defined as follows:

3.1 Basic Principles of Building Response

Rv 5 Ra 5

ωe Rd ω ω 2 e

ω

(3.18) Rd

(3.19)

In order to analyze the system response, the variation in the dynamic response factors and phase angle is presented (Fig. 3.4) as a function of the damping ratio and the system frequency ratio, ω e : ω e 5 ωe =ω

(3.20)

When ω e is equal to 1, the system is at resonance. In systems with no damping (ζ 5 0Þ, this means the amplitude of the response goes to infinity. Damping prevents this behavior by limiting the maximum amplitude, as described by Eq. (3.16). Damping also shifts the forcing frequency at which the maximum displacement amplitude is attained. As can be seen in Fig. 3.4, for lightly damped

FIGURE 3.4 Deformation response factor (Rd) and phase angle (ϕ) as a function of frequency ratio for a damped system excited by a harmonic force.

45

46

CHAPTER 3 Damping considerations in tall buildings

FIGURE 3.5 Arbitrary force represented by several infinitesimal impulses. Adapted from Chopra, A.K., 2012. Dynamics of Structures: Theory and Application to Earthquake Engineering, Prentice Hall, Upper Saddle River, New Jersey.

systems (ζ , 10%Þ, the peak response occurs at ω e  1, but as damping increases, the peak displacement occurs at lower values of ω e . Where the SDOF system is excited by an arbitrary force like a seismic force, then F ðtÞ 5 2 mu¨g ðtÞ, where u¨g ðtÞ represents earthquake ground acceleration. To solve the equation of motion, this force can be characterized as a sequence of infinitesimally short impulses (Fig. 3.5). Consequently, the solution of Eq. (3.5) leads to the following expression: uðtÞ 5

ðt

F ðτ Þhðt 2 τ Þ dτ

(3.21)

0

where hðt 2 τ Þ is the unit impulseresponse function of a damped SDOF system. Considering the excitation as a unit impulse ground motion, the system response is given by Duhamel’s integral as follows: uðtÞ 5

1 mωD

ðt

F ðτ Þe2ζωðt2τ Þ sin½ωD ðt 2 τ Þ dτ

(3.22)

0

This integral is useful in obtaining the response of a linear SDOF system under arbitrary forces. If F ðτ Þ is a simple function, the closed-form solution of the integral is possible. However, in most stochastic events (wind or earthquake event), solving the integral needs more complicated numerical methods (e.g., time-stepping methods). The equation of motion of a damped SDOF system when subjected to an earthquake can be expressed by substituting F ðtÞ 5 2mu¨g ðtÞ in Eq. (3.5) and dividing by the system mass m as: u¨ðtÞ 1 2ζωu_ ðtÞ 1 ω2 uðtÞ 5 2 u¨g ðtÞ

(3.23)

The deformation response of the system depends on two parameters: the system frequency ω (or period T) and the damping ratio ζ. To efficiently solve the

3.1 Basic Principles of Building Response

FIGURE 3.6 An idealized three DOF system.

above equation, numerical procedures such as Newmark’s method can be used (see Chopra (2012) for further explanations).

3.1.2 MULTIPLE DEGREES-OF-FREEDOM SYSTEM When more than one independent coordinates (i.e., more than one displacement modes) are required to express the overall motion of a system, an MDOF system is needed to idealize the structure (Fig. 3.6).

3.1.2.1 Equations of motion Similar to the SDOF system, the equation of motion of an MDOF system can be developed on the basis of the equilibrium of effective forces corresponding to each degree of freedom (DOF): FIi ðtÞ 1 Fdi ðtÞ 1 FSi ðtÞ 5 Fi ðtÞ

(3.24)

where FIi ðtÞ is the inertial force associated with the ith DOF, Fdi ðtÞ the damping force associated with the ith DOF, FSi ðtÞ the spring force associated with the ith DOF, and Fi ðtÞ the applied dynamic force associated with the ith DOF (due to wind or earthquake). Assuming that the MDOF system has N degrees of freedom, the inertial force, damping force, and spring force associated with the ith DOF are expressed as follows, respectively: FIi ðtÞ 5 mi1  u¨1 ðtÞ 1 mi2  u¨2 ðtÞ 1 . . . 1 miN  u¨N ðtÞ

(3.25)

Fdi ðtÞ 5 ci1  u_ 1 ðtÞ 1 ci2  u_ 2 ðtÞ 1 . . . 1 ciN  u_ N ðtÞ

(3.26)

FSi ðtÞ 5 ki1  u1 ðtÞ 1 ki2  u2 ðtÞ 1 . . . 1 kiN  uN ðtÞ

(3.27)

where i 5 1 to N (i.e., a set of N equations is obtained for each of the above forces) and mij mass associated with the ith DOF due to unit acceleration along the jth DOF (i.e., mass influence coefficient) cij damping constant associated with the ith DOF due to unit velocity along the jth DOF (i.e., damping influence coefficient)

47

48

CHAPTER 3 Damping considerations in tall buildings

kij stiffness associated with the ith DOF due to unit displacement along the jth DOF (i.e., stiffness influence coefficient) Substituting Eqs. (3.25)(3.27) in Eq. (3.24), the equation of motion in matrix form can be expressed as: M u¨ðtÞ 1 CU_ ðtÞ 1 KU ðtÞ 5 FðtÞ

(3.28)

where M is the mass matrix (N 3 N), C the damping matrix (N 3 N), K the stiffness matrix (N 3 N), FðtÞ the force vector (N 3 1), U¨ ðtÞ the acceleration vector (N 3 1), U_ ðtÞ the velocity vector (N 3 1), and U ðtÞ the displacement vector (N 3 1). The format of above matrices and vectors is as follows, respectively: 2

3 m11 ? m1N M54 ^ & ^ 5 mN1 ? mNN 2 3 c11 ? c1N C54 ^ & ^ 5 cN1 ? cNN 2 3 k11 ? k1N K54 ^ & ^ 5 kN1 ? kNN 2 3 F1 ðtÞ FðtÞ 5 4 ^ 5 FN ðtÞ 2 3 u¨1 ðtÞ U¨ ðtÞ 5 4 ^ 5 u¨N ðtÞ 2 3 u_ 1 ðtÞ U_ ðtÞ 5 4 ^ 5 u_ N ðtÞ 2 3 u1 ðtÞ 4 U ðtÞ 5 ^ 5 uN ðtÞ

(3.29)

(3.30)

(3.31)

(3.32)

(3.33)

(3.34)

(3.35)

3.1.2.2 Free vibration properties For an MDOF system with N DOF, a series of N independent (displacement) vectors, φ, can be identified. These are called natural modes of vibration or mode shapes. Fig. 3.7 illustrates the mode shapes of an idealized two-story building structure with two DOF (Only horizontal displacements are considered).

3.1 Basic Principles of Building Response

FIGURE 3.7 Natural vibration modes of an idealized two DOF building system.

Consider the undamped, free-vibration case, for example, the damping matrix is zero (i.e., C 5 0), and the external force vector is a zero vector (i.e., FðtÞ 5 0) in Eq. (3.28), undamped, free vibration of the MDOF system can be expressed by: M U¨ ðtÞ 1 KU ðtÞ 5 0

(3.36)

The response of the structure can be modeled by the multiplication of two independent variables: the set of mode shape vectors φm ðxÞ that control the shape of the deformation at location x, and a scalar function of time qm ðtÞ that controls the amplitude of the motion. Mathematically, this can be written as: U ðtÞ 5 φqðtÞ

(3.37)

where φ is a matrix containing all the mode shapes, and qðtÞ is a vector containing all the individual functions qm ðtÞ. It can be shown (Chopra, 2002) that the solution of Eq. (3.36) requires qm ðtÞ to be harmonic; therefore, substituting a harmonic function of circular frequency ωm for the time-dependent displacement [qm ðtÞ 5 Am cosωm t 1 Bm sinωm t] in Eq. (3.37), and then the resulting expression in Eq. (3.36), the following equation is obtained after mathematical manipulations: 

 2 ω2m Mφm 1 Kφm qm ðtÞ 5 0

(3.38)

The constants Am and Bm can be obtained using the initial conditions of displacement and velocity. Since qm ðtÞ 5 0 is its trivial solution, the solution of following real eigenvalue problem gives the vibration properties of the MDOF system: 

 K 2 ω2m M φm 5 0

(3.39)

Neglecting the trivial solution (φm 5 0), the nontrivial solution of the above equation reads:

49

50

CHAPTER 3 Damping considerations in tall buildings det K 2 ω2m M 5 0

(3.40)

which results in N real and positive frequencies ωm (m 5 1; 2; . . . ; NÞ with the arrangement ordered as ω1 , ω2 , . . . , ωN . Having determined the frequencies ωm , the solution of Eq. (3.39) gives the corresponding vectors φm (mode shapes). Note that the solution of Eq. (3.39) for φm represents the shape of a vector with relative values (not absolute values) of N displacements (degrees of freedom): 8 9 φ1m > > > > > > φ2m > > < = φm 5 φ3m > > > > ^ > > > > : ; φNm

(3.41)

Assembling all the N frequencies and corresponding Nm modes into a compact matrix form, the spectral (diagonal) matrix of eigenvalue problem Ω2 and the modal matrix Φ are expressed, respectively, by: 2

ω21 0 6 0 ω2 2 2 Ω 56 4 ^ ^ 0 0 2 φ11 φ12 6φ φ 6 Φ 5 4 21 22 ^ ^ φN1 φN2

3 ... 0 7 ... 0 7 ^ 5 & . . . ω2 N 3 . . . φ1Nm 7 . . . φ2Nm 7 5 ^ & . . . φNN m

(3.42)

(3.43)

An important property of the natural modes is the orthogonality of different modes, that is, ωm 6¼ ωr , as expressed below: φTm Kφr 5 0 φTm Mφr 5 0

(3.44)

φTm

where is the transpose of the mth mode shape vector. The orthogonality condition demonstrates that the following matrices are diagonal: K 5ΦT KΦ M 5ΦT MΦ

(3.45)

where the diagonal elements (generalized, modal, stiffness, and mass) read: K m 5 φTm Kφm 5 ω2m M m

M m 5 φTm Mφm

(3.46)

The normalization of natural modes is an important process to standardize the elements associated with different DOF. There are several approaches to normalize each mode: • • •

Normalize by the amplitude of the DOF with the largest modal amplitude Normalize by the modal amplitude at the roof Normalize to achieve M m 5 φTm Mφm 5 1 (computer programs commonly apply this approach)

3.1 Basic Principles of Building Response

In addition, for earthquake engineering purposes, it is useful to define the mode participation factor as follows: Γm 5

φTm Mr Mm

(3.47)

where r is the influence vector which takes into account how the ground acceleration is transmitted to the different DOF. If the ground motion is purely horizontal (no rotational input at the base), and only horizontal DOF are considered, then r becomes a vector of 1s and induces a rigid body motion in all modes. It can be shown that the product of the mode participation factor and the mode shape of the mth mode, Γm φm , is independent of how the modes are normalized. Therefore it can be a better tool to understanding the behavior of the building than the mode shape alone as it gives insight into its lateral deflected shape. Fig. 3.8 shows the product Γm φm for a generic shear building. It can be seen, for example, how the first and second modes have opposite signs at the roof, and therefore, their sum diminishes the total response of the building at that level. This product is sometimes referred to as the effective mode shape of the building, and it is primarily affected by the lateral load-resisting system of the building and the distribution of stiffness along its height (Miranda and Akkar, 2005). For a damped MDOF system, free vibration leads to the following equation of motion [by substituting FðtÞ 5 0 into Eq. (3.28)]: M U¨ ðtÞ 1 CU_ ðtÞ 1 KU ðtÞ 5 0

(3.48)

Solutions to the above equation may differ depending on whether the format of the damping matrix is classical or nonclassical. The damping matrix is classical if the following condition is met (Caughey and O’Kelly, 1965):

FIGURE 3.8 Effective mode shapes for the first four modes of a generic shear building. Adapted from Cruz, C., 2017. Evaluation of Damping Ratios Inferred from the Seismic Response of Buildings. PhD Thesis Stanford University, California.

51

52

CHAPTER 3 Damping considerations in tall buildings

CM 21 K 5 KM 21 C

(3.49)

In this case, all the natural modes of the system are identical to those of the undamped system. For a classically damped MDOF system, all the modes are uncoupled, and thus premultiplying Eq. (3.48) by ΦT , the following equation of motion is obtained: M q€ ðtÞ 1 C q_ ðtÞ 1 KqðtÞ 5 0

(3.50)

where M and K are defined in Eq. (3.45), and the damping matrix is diagonal: C 5ΦT CΦ

(3.51)

Eq. (3.50) gives N uncoupled differential equations in modal coordinates qm ðtÞ for a classically damped system, as follows: M m q€ m ðtÞ 1 C m q_ m ðtÞ 1 K m qm ðtÞ 5 0

(3.52)

where M m and K m are expressed in Eq. (3.46) and the generalized (modal) damping is given by: C m 5 φTm Cφm

(3.53)

Classical modal analysis is valid for a classically damped MDOF system. Indeed, for each mode (e.g., mth mode), a damping ratio can be derived in the way similar to the case of the SDOF system in Section (3.1.1), which is expressed as: ζm 5

Cm 2M m ωm

(3.54)

Dividing Eq. (3.52) by M m , the following equation in a modal form is found: q€ m ðtÞ 1 2ζ m ωm q_ m ðtÞ 1 ω2m qm ðtÞ 5 0

(3.55)

The solution of the above equation is identical to that of Eq. (3.6) of an SDOF system and reads:

 q_ ð0Þ 1 ζ m ωm qm ð0Þ qm ðtÞ 5 e2ζ m ωm t qm ð0Þ cosωmD t 1 m sinωmD t ωmD

(3.56)

where the damped frequency corresponding to the mth mode is defined by: ωDm 5 ωm

qffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 ζ 2m

(3.57)

If the damping matrix does not satisfy Eq. (3.49), then the system is considered to have nonclassical damping. In this case, the natural modes are complex (nonreal values), and the damping matrix C is not diagonal. To solve the equation of motion for a nonclassically damped system, other procedures, for complex eigenvalue analysis, are employed. This is beyond the scope of this book, and interested readers should refer to Chopra (2012).

3.1 Basic Principles of Building Response

3.1.2.3 Forced vibration properties The equation of motion of a damped MDOF system (Eq. (3.28)) can be solved in the similar way as an SDOF system mentioned above. If the MDOF system with N DOF is excited by both an external force to the main body, FðtÞ, and a ground acceleration force to the base, 2 M1u€ g ðtÞ, the equation of motion of the system reads: M U€ ðtÞ 1 CU_ ðtÞ 1 KU ðtÞ 5 FðtÞ 2 M1u€ g ðtÞ

(3.58)

where 1 is called the effective vector of order N with elements equal to unity. Note that this equation is valid only if the building has a fixed condition at the base, that is, the flexibility of the soil is neglected (typical assumption). The above matrix-form equation contains N differential equations, each of which corresponds to a particular DOF. Therefore the relative displacement associated with each DOF can be computed using numerical methods, similar to those applied for the SDOF system in Section 3.1.1.3. Note that the total acceleration of the system is represented by the combination of relative acceleration of the main body (building stories) and the ground acceleration u€ g ðtÞ as follows (Fig. 3.9): t U€ ðtÞ 5 u€ g ðtÞ1 1 U€ ðtÞ

(3.59)

3.1.3 EQUIVALENT VISCOUS DAMPING The solution to the dynamic problem of a nonlinear structural system can be simplified utilizing an equivalent value of viscous damping. This parameter also allows for the combined effect of different dissipation mechanisms in a structure using a single damping coefficient. It is important to note that this is an approximation meant to simplify the mathematical equations governing the system, but this simplification has trade-offs. A good summary of these trade-offs can be found in Priestley (1993) and Miranda (2006). The equivalent damping (ζ eq ) is derived considering one cycle of resonant steady-state harmonic oscillation (i.e., excitation frequency is equal to system frequency), which is defined as the ratio between the energy dissipated by the system per cycle (ED ) and the available strain energy (ES0 ), as follows (Fig. 3.10): 1 ED 4π ES0

(3.60)

1 ES0 5 ku20 2

(3.61)

ζ eq 5

with

where k is the system stiffness and u0 is the amplitude of harmonic motion.

53

54

CHAPTER 3 Damping considerations in tall buildings

FIGURE 3.9 Schematic accelerations in a (A) building frame and (B) tower-like system.

FIGURE 3.10 Parameters needed to calculate equivalent damping ratio for viscous damping.

In the case of a nonresonant harmonic loading, other damping models are more appropriate, such as a hysteretic damping model (as introduced by Clough and Penzien (2003)). The concept of equivalent damping ratio for hysteretic damping models is described in Section 4.1.1.2.

3.2 Damping in Tall Buildings

3.2 DAMPING IN TALL BUILDINGS In any building, damping can be decomposed into four different contributions: intrinsic/inherent (or structural) (ζ i ), aerodynamic (ζ a ), hysteretic (ζ h ), and supplementary/additional (ζ d ). The total damping (ζ T ) is then defined by adding the different damping matrices. Intrinsic damping is usually estimated from empirical formulas based on databases of full-scale measurements. Aerodynamic damping is related to the viscosity of the air and the speed of the object relative to the airflow (Kareem and Gurley, 1996). This type of damping is typically neglected for tall buildings (Marukawa et al., 1996), but it could become relevant in the near future for the designs of flexible supertall buildings. Hysteretic damping is the energy dissipation resulting from the inelastic behavior of structural elements in cases of severe loading conditions (e.g., earthquake). Supplementary damping comes from specific devices added to the building system. In general, available estimation methods only consider the damping of the fundamental mode. However, higher modes are particularly important for tall buildings, and damping due to these higher modes may impact the building serviceability (Kareem and Gurley, 1996). An estimation of higher mode damping was developed by Kareem (1981) based on the limited data available at the time, proposing a stiffness-proportional equation for estimating modal damping. Recently, Cruz and Miranda (2017a) carefully examined the relationship between modal damping ratios and their corresponding modal frequencies of 24 buildings subjected to different earthquakes motions, proposing the following equation for the damping ratio of higher translational modes:



 fm ζ ðf Þ 5 ζ 1 1 1 C 21 f1

(3.62)

where C is a constant equal to 0.12, and fm and f1 represent the natural frequency for the mth and first mode, respectively. It has been argued that this linear variation of modal damping with frequency is a consequence of soilstructure interaction (Cruz and Miranda, 2017b). This is an important progress in design of tall buildings, but a great deal of research still needs to be conducted in this field in order to obtain more accurate damping estimates for higher modes.

3.2.1 DAMPING MODELING The inherent damping matrix is usually defined by two general approaches: 1. Rayleigh damping:

55

56

CHAPTER 3 Damping considerations in tall buildings

The damping matrix and the fraction of total damping ζ T for the mth mode of vibration with modal period of Tm are, respectively: C 5 aM M 1 aK K ζ Tm 5

aM Tm aK π 1 Tm 4π

(3.63) (3.64)

where aM and aK are the proportionality constants for mass and stiffness matrices, respectively, determined based on two periods of vibration, Ti and Tj (corresponding to ith and jth modes): aM 5 4πζ T  aK 5

1  Ti 1 Tj

ζ T Ti Tj   π Ti 1 Tj

(3.65) (3.66)

In nonlinear analysis, the stiffness matrix does not remain constant (the structure usually softens); therefore the damping matrix also varies, in contrast to an elastic analysis where both remain unchanged. Moreover the mass-proportional component can lead to unrealistic values for large rigid body motions (Hall, 2005). For these reasons, Charney (2006) proposed three different damping matrix definitions for a nonlinear system: • Based on the initial stiffness (i.e., constant) • Based on constantly updating (throughout the analysis) the tangent stiffness with the same initial proportionality constants, aM and aK • Based on constantly updating (throughout the analysis) the tangent stiffness with updated proportionality constants, aM and aK , to maintain a specified critical damping percentage for the inelastic vibration modes Despite the numerous research investigations conducted, there is no consensus on which option is optimal to utilize (Charney, 2006; Petrini et al., 2008). It is also important to mention that Rayleigh damping is employed because it simplifies the equations of motion while producing a classical damping matrix. In general the evidence in the literature does not support mass-proportional damping. A discussion of these issues can be found in Cruz and Miranda (2017a). 2. Modal damping (combination of elastic vibration modes):  21 C 5 ΦT Ci Φ21

(3.67)

where Ci is the diagonal matrix of damping coefficients with the mth diagonal element of 2ζ m M m ωm . Here M m is defined in Eq. (3.46).

3.2 Damping in Tall Buildings

Subsequently the previous equation can be implemented as follows: C5M

N X 4πζ Tm m51 Tm M m

! φm φTm

M

(3.68)

where M is a diagonal mass matrix and N is the total number of modes included. Modal damping can be used to specify the damping matrix to be used in an inelastic analysis to avoid the apparition of spurious damping forces (Chopra and McKenna, 2016).

3.2.2 INTRINSIC DAMPING While calculating the total damping for the system, it is necessary to estimate the intrinsic damping of the structural system (i.e., the damping from the building without supplementary systems). This has been the subject of many researches, and while there are many measurements of damping under low wind speeds, there are very few under-the-wind speeds considered for service conditions (i.e., the 1-year wind), let alone ultimate wind speeds or seismic effects. In the following section, justifications for the estimation of intrinsic damping are reviewed in detail.

3.2.2.1 Intrinsic damping sources Intrinsic damping estimation is not trivial from a theoretical point of view. For this reason, field measurements have been extensively used in the past for estimating intrinsic damping from buildings behavior when excited by wind and/or seismic forces (e.g., Davenport and Hill-Carroll, 1986; Ellis, 1996; KijewskiCorrea and Pirnia, 2007; Lagomarsino and Pagnini, 1995; Satake et al., 2003; Yoon and Ju, 2004). Several databases have been collected worldwide and one of the first estimations of intrinsic damping ratio refers to case study of five highrise buildings (2030 stories high) excited by underground nuclear explosions (Blume, 1970). Subsequently a wider database of intrinsic damping ratios was collected from 74 buildings (Scholl, 1975), 139 buildings (Haviland, 1976), and 165 buildings (Davenport and Hill-Carroll, 1986), mainly under low-amplitude motions (e.g., forced vibration, ambient vibration, wind, etc.). This database was extended by Lagomarsino (1993) with 185 buildings again subjected to lowamplitude motions. A larger database was collected by Fritz et al. (2009) in which 1572 damping ratios were measured using full-scale tests on 972 buildings around the world (271 data points from 98 buildings subjected to wind or ambient forces; 260 points from 93 buildings subjected to earthquake; and 771 data points from 273 buildings under forced vibration tests). The majority of these data come from the measurements in Japan and the United States. The latest projects in this research area are the ongoing research project at the University of Notre Dame concerning three buildings in Chicago (Kijewski-Correa et al., 2006) and a number of buildings currently being monitored in China (Guo et al., 2012; Li et al., 1998, 2003, 2004, 2008, 2011).

57

58

CHAPTER 3 Damping considerations in tall buildings

One of the major databases for intrinsic damping for the fundamental mode of tall buildings is given by Satake et al. (2003). This was reviewed and updated by Smith et al. (2010) that subsequently analyzed the data (Fig. 3.11) with the following considerations: •

•

•

• •

There is a lower “floor” of around 0.3% of critical intrinsic damping. This might be interpreted as representing the intrinsic damping of the structural materials. There is significant scatter in the data. This may be the result of variability in the type and age of the buildings, type of nonstructural fit-out, foundation type, amplitude of excitation, etc. There is a distinct downward trend in the average value of damping as building height increases. While there is still a high degree of scatter for taller buildings, there are virtually no measurements above 1% in buildings over 250 m high. Based on simple observations, there is no clear correlation between the level of damping and material type. An outlier measurement of 1.5% at 318 m was the reading in Allied Bank Plaza measured during Hurricane Alicia (Halvorson and Isyumov, 1986). The wind speed measured corresponded to an unfactored 50-year wind.

The major reason for the scatter in the data comes from the numerous intrinsic damping sources, among which the most important are considered to be (Tamura, 2012): • • • • • • • •

Material damping Friction between members and connections Structural system and joint types Foundation and soil types (soilstructure interaction) Interior partitions Exterior cladding Other nonstructural members Vibration amplitude

These intrinsic damping sources follow different laws. For example, material damping can be modeled with hysteretic behavior whereas friction is usually modeled with Coulomb’s law. Among all these intrinsic damping sources, the most relevant ones for tall buildings are soilstructure interaction (Cruz and Miranda, 2017b) and structural behavior (“frame-action” or “cantilever/axial shortening action”) (Figs. 3.12 and 3.13) (Bentz and Kijewski-Correa, 2008; Bentz, 2012; Erwin et al., 2007; Kijewski-Correa et al., 2006). In the case of frame action (Fig. 3.12), the building is behaving like a moment frame system in which the lateral loads induce bending deformations in the elements. Therefore it is expected to produce larger damping due to shear and flexural behavior of frame members and flexibility of the connections (Bentz and Kijewski-Correa, 2008;

(A) 10

Trendline Damping = 7⋅20h–0⋅3887 R2 = 0⋅1866

Damping rao: %

Damping rao: %

(B) 10

Buildings, steel

9 8 7 6 5 4 3 2

Damping = 21⋅2h–0.5708 R2 = 0⋅3161

1 0

1 0 0

100

200 300 Height h: m

400

500

(C)

0

100

200 300 Height h: m

400

(D) 10 10 9 8 7 6 5 4 3 2 1 0

Buildings, reinforced concrete Damping rao: %

Damping rao: %

Buildings, steelreinforced concrete Trendline

9 8 7 6 5 4 3 2

Trendline Damping = 28⋅25h–0⋅6393 R2 = 0⋅4642

0

100

200 300 Height h: m

400

500

9 8 7 6 5 4 3 2 1 0

Chimneys, reinforced concrete Trendline Damping = 0⋅0021h1.0506

0

100 200 Height h: m

300

FIGURE 3.11 Primary source of data for measurements of fundamental mode intrinsic damping in high-rise buildings (Smith et al., 2010).

500

60

CHAPTER 3 Damping considerations in tall buildings

FIGURE 3.12 Frame action for damping estimation comparison. Adapted from Van Den Berg, R.L.J., 2012. Investigation of Damping in High-Rise Buildings. Graduation Project, Tu Delft, Netherland.

FIGURE 3.13 Cantilever action for damping estimation comparison. Adapted from Van Den Berg, R.L.J., 2012. Investigation of Damping in High-Rise Buildings. Graduation Project, Tu Delft, Netherland.

Bentz, 2012). In the case of cantilever/axial shortening (Fig. 3.13), the building behaves like a brace frame system in which the lateral loads induce axial deformations in the elements. Therefore the dissipation comes from axial shortening (Bentz and Kijewski-Correa, 2008; Bentz, 2012).

3.2 Damping in Tall Buildings

FIGURE 3.14 Stick-slip model for intrinsic damping in buildings.

3.2.2.2 Intrinsic damping estimation Intrinsic damping energy losses were theoretically explained by Wyatt (1977) with a friction model called “stiction.” In this model the contact surfaces begin to slip at certain amplitudes, thus losing their stiffness. Moreover the number of slipping surfaces increases with amplitude. This is defined as “stick-slip” behavior by Davenport and Hill-Carroll (1986) (Fig. 3.14). This stick-slip behavior happens at two different scales: between structural elements and at the material micro-scale level (Jeary, 1986). As can be seen, the mechanism behind intrinsic damping is mainly friction behavior. However, Wyatt (1977) demonstrated that a series of friction mechanisms integrated over a range of amplitudes produced a response similar to a viscous-like system. One of the most interesting aspects of dynamic properties of tall buildings is their sensitivity to vibration amplitude. Intrinsic damping is mainly nonlinearly increasing with amplitude while natural frequency decreases (Satake et al., 2003). This damping increase with amplitude is well correlated with Jeary’s theoretical model (Fig. 3.15) (Jeary, 1986, 1996), in which, at higher amplitudes, more and smaller cracks will form in the structural elements such as concrete members and items like slip-critical bolted connections for steel members. However, no consensus is found in the literature as damping can decrease with amplitude (Smith et al., 2010) or decrease after a “critical tip deflection” that is very small (on the order of 1025 to 1024 of the building height) (Tamura, 2012). This “critical tip deflection” does not agree with Jeary’s theoretical model that predicts a drop in damping in the highest amplitude range. Given this complex behavior and the different sources of intrinsic damping (as shown in the previous section), full-scale measurements are the best means to

61

62

CHAPTER 3 Damping considerations in tall buildings

FIGURE 3.15 Jeary’s theoretical intrinsic damping model.

estimate modal properties for a building. However, collecting full-scale data is just one part of the process since appropriate damping estimation techniques must be used. Several different methods have been summarized by Tamura (2012) and Cruz and Miranda (2016). When identifying the building’s dynamic properties from seismic records, most of the system identification methods currently employed use both input and output signals (Cruz and Miranda, 2016). However, when estimating damping from ambient or wind-induced vibrations, the most suitable ones are based on the “Output-Only Modal Estimation” (OOME) because it is not possible to accurately characterize the forcing function (Brincker et al., 2000). Other traditional damping estimation techniques can be used, but they are based on known input: free decay damped response (FDDR), frequency response function (FRF), and impulse response function (IRF). The main sources of errors and discrepancies in these damping estimation techniques are due to factors (Tamura, 2012) such as: • • •

Modal properties evaluation techniques Damping evaluation methods Degree of nonstationary of excitations

It is clear that given the available estimation theories, obtaining consistent damping values is difficult. Indeed, due to broadband excitation, the damping information encoded in the response data, called “Fisher” information (Bernal et al., 2012), is low. Consequently this leads to a high variability in the damping estimation, even for the same structure when different data sets are used (Bernal et al., 2012).

3.2.2.3 Intrinsic damping prediction In literature there are several prediction formulas available for calculating intrinsic damping based on the analysis of different building databases. There are two major typologies: (1) power law models [e.g., Davenport and Hill-Carroll (1986))

3.2 Damping in Tall Buildings

and (2) piecewise linear models (e.g., Lagomarsino (1993)]. The first one is based on the assumption that the damping is mainly caused by friction. The second model is based on the definition of three amplitude-dependent regions: low plateau, high plateau, and a nonlinear transition zone. However, as already explained (Section 3.2.2.2), in some cases, increase in amplitude may decrease intrinsic damping (Tamura, 2005) due to an amplitude- and frequency-dependent correlation (Kijewski-Correa et al., 2007). Therefore for this specific case, there are no general prediction models (Spence and Kareem, 2014). For wind-resistant analyses, a summary of the major predictions is given in the following: •

Jeary (1986), based on pffiffiffiffi L1 =2

ζ i1 5 0:01f1 1 10

•

α1 α2 x H α2

x α 1 βf1 1 γ f1 H

(3.71)

where α, β; and γ are parameters related to the building characteristics. Tamura et al. (2000) ζ i1 5 γ 1 f1 1 γ 2 1

•

(3.70)

where α1 and α2 are parameters depending on the number of stories and primary construction material. Lagomarsino (1993) ζ i1 5

•

(3.69)

where L1 is the building dimension along vibration direction (m), x is the tip displacement amplitude of vibration for the cycle being evaluated, H is the building height, and f1 is the first mode natural frequency. Davenport and Hill-Carroll (1986) ζ i1 5

•

x=H

γ3 x H

(3.72)

where γ 1 , γ 2 ; and γ 3 are parameters depending on primary construction material. Satake et al. (2003) and Tamura (2012) Amplitude-independent formula ζ i1 5 0:013f1 ðsteelÞ ζ i1 5 0:014f1 ðRCÞ

(3.73)

Amplitude-dependent formula if x=H # 2 3 1025 x 2 0:0018 ðsteelÞ for H # 200 m H x ζ i1 5 0:013f1 1 470 1 0:0029 ðRCÞ for H # 130 m H

ζ i1 5 0:014f1 1 470

(3.74)

63

64

CHAPTER 3 Damping considerations in tall buildings

For higher modes, damping ratios with m 5 2 and 3:

•

ζ im 5 ð1:3  1:4Þζ i;m21 ðsteelÞ ζ im 5 1:4ζ i;m21 ðRCÞ ζ im 5 ð1:7  1:8Þζ i;m21 ðSRCÞ

(3.75)

ζ i1 5 1:945 1 0:195f13:779

(3.76)

Zhang and Cho (2009)

It can be seen in all the above expressions that there is a damping dependency on frequency while in some others there is also dependency on the amplitude. Each damping model results in a different damping prediction, and Figs. 3.16 and 3.17 show a comparison between them (Aquino and Tamura, 2013c; Spence and Kareem, 2014). As results are not uniform a common consensus has not been reached on which prediction model to utilize. Concerning seismic-resistant analyses, although no uniform consensus exists, it is possible to use measured intrinsic damping from available databases such as Celebi et al. (1993) and Celebi (1996). In addition, according to some more recent studies (Bernal et al., 2015; Cruz and Miranda, 2016; PEER/ATC, 2010; PEER, 2017), inherent damping in tall buildings under low-to-moderate (seismic) amplitudes is lower than that in low-rise buildings, mainly due to the limited damping influence from foundations in tall buildings. However, for the sake of an approximate estimation, the following damping predictors may be suggested: •

Fritz et al. (2009)

FIGURE 3.16 Comparison of different prediction models with building height.

3.2 Damping in Tall Buildings

FIGURE 3.17 Modal frequency against modal damping ratios for the first three modes (A) steel and (B) RC (Spence and Kareem, 2014).

0:06 ffiffiffiffi ðsteelÞ ζ i1 5 p 4 N 0:14 ζ i1 5 pffiffiffiffi ðRCÞ N

(3.77)

0:09 ζ i1 5 pffiffiffiffi ðSRCÞ N

where N is the number of building stories (up to 100). Note that Eqs. (3.77) may be considered somehow as general predictors (for both wind and seismic analyses) since the data source includes both excitation types. •

Bernal et al. (2015) ζ i1 5 1:2 1 4:26e20:013H ðsteelÞ ζ i1 5 3:01 1 3:45e20:019H ðRCÞ

•

(3.78)

Cruz and Miranda (2016) ζ i1 5 0:9H 20:76 ðsteelÞ ζ i1 5 15:6H 21:31 ðRCÞ

(3.79)

A comprehensive study that paid special attention to the reliability of the inferred damping values can be found in Cruz (2017). In this work the following formula is recommended: ζ i1 5 0:21H 20:47 for H , 50 m ζ i1 5 0:51H 20:68 for H . 50 m

(3.80)

65

66

CHAPTER 3 Damping considerations in tall buildings

Moreover Cruz (2017) found that material (steel or RC) was not statistically significant in the damping ratio of buildings subjected to earthquakes and that the amplitude of the response had a negligible effect after a critical/threshold amplitude was reached. According to damping values inferred from 119 seismic responses in 24 buildings in California, Cruz and Miranda (2017a,b) concluded that higher modes damping ratios increase approximately linearly with increasing modal frequency.

3.2.2.4 Intrinsic damping code limitations Previous sections have shown that there is no consensus among researchers in the estimation of the intrinsic damping value for tall buildings. This great variability is also reflected in code and guideline recommendations, since they do not provide prescriptive theoretical models but only recommend values for structural analyses (that are, in most of the case, valid for low-rise buildings). In any case, it is important to understand that, regardless of the code or standard utilized, building damping sensitivity analysis should be carried out by designers to understand the damping influence on the building response. In the following, a general overview of the current codes/guidelines for damping recommendations is given: •

• •

Australia/New Zealand: AS/NZS 1170.2.2011 Part 2 (Wind Actions) (AS/NZS, 2011) and AS/NZS 1170.5.2004 Supp 1 (Earthquake Actions) (AS/NZS, 2004) These codes state that users should seek adequate/sources for the estimation of intrinsic damping that take into account building height and amplitude of vibration. However, for wind analyses the following limitations are given (AS/NZS, 2011): • Ultimate limit state intrinsic damping: • Steel structures: 2% • RC structures: 3% • Serviceability limit state intrinsic damping: • Steel structures: 1.2% for deflection calculations and 1% for acceleration calculations at the top of buildings and towers • RC structures: 1.5% for deflection calculations and 1% for acceleration calculations at the top of buildings and towers Instead, for earthquake analyses 5% for all modes or as specified by the appropriate material standard. Eurocode 1 (CEN, 2010): Provides in an appendix, the logarithmic decrement of intrinsic damping for wind action at the fundamental mode for several different structural categories, as shown in Table 3.1. The logarithmic decrement can be associated with damping with the following expression:

3.2 Damping in Tall Buildings

Table 3.1 Eurocode 1 (CEN, 2010) Intrinsic Damping Recommendations

Structural Type Reinforced concrete buildings Steel buildings Mixed structures concrete 1 steel Reinforced concrete towers and chimneys Unlined welded steel stacks without external thermal insulation Unlined welded steel stacks with external thermal insulation Steel stack with one liner with external thermal insulation

Steel stack with two or more liners with external thermal insulation

h/ba # 18 20 # h/ba # 24 h/ba $ 26 h/ba # 18 20 # h/ba # 24 h/ba $ 26

Steel stack with internal brick liner Steel stack with internal gunite Coupled stacks without liner Guyed steel stack without liner Steel bridges 1 lattice steel towers

Welded Highresistance bolts Ordinary bolts

Composite bridges Concrete bridges

Prestressed without cracks With cracks

Timber bridges Bridges: aluminum alloys Bridges: glass or fiber reinforced plastic Cables

a

Parallel cables Spiral cables

h is the building height and b the building base.

Logarithmic Decrement Structural Damping ðLδ Þ

Intrinsic Damping   ζ i (%)

0.1 0.05 0.08 0.03 0.012

1.6 0.8 1.3 0.5 0.2

0.02

0.3

0.02 0.04

0.3 0.6

0.014 0.02 0.04

0.2 0.3 0.6

0.025

0.4

0.07 0.03 0.015 0.04

1.1 0.5 0.2 0.6

0.02 0.03

0.3 0.5

0.05

0.8

0.04

0.6

0.04

0.6

0.1

1.6

0.06 0.12 0.02 0.04 0.08 0.006

1.0 1.9 0.3 0.6 1.3 0.1

0.02

0.3

67

68

CHAPTER 3 Damping considerations in tall buildings

1 ζ i1 5 rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  ffi 11

•

•

•

•

•

2π Lδ

2

(3.81)

where Lδ is the factor accounting for logarithmic decrement of structural damping. Germany DIN1055 Teil 4 (DIN, 2005): The wind code provides recommended values of logarithmic decrement, similar to those of Eurocode 1 (CEN, 2010). • Steel and aluminum structures:  Bolted connection Lδ 5 0:05 - ζ i1 5 0:5%0:8%  Welded connection Lδ 5 0:02 - ζ i1 5 0:3%  Supplement for absorbing internals, for example, linings Lδ 5 0:02 ζ i1 5 0:3%  Supplement for open-bolted lattice structures Lδ 5 0:02 - ζ i1 5 0:3% • Concrete and reinforced concrete constructions:  Without cracks Lδ 5 0:04 - ζ i1 5 0:6%  With cracks Lδ 5 0:10 - ζ i1 5 1:6%  Prestressed Lδ 5 0:04 - ζ i1 5 0:6%  Supplement for absorbing internals, for example, linings Lδ 5 0:02 ζ i1 5 0:3% • Masonry structures Lδ 5 0:12 - ζ i1 5 1:9% • Wood structures Lδ 5 0:15 - ζ i1 5 2:4% ISO/CD 3010 (ES ISO, 2017): This code states that for wind design, a damping ratio of 1% is sufficient while for seismic design a higher damping ratio is necessary due to the larger deformations involved. For example, for regular structures, the first mode damping ratio range is between 2% and 5%, depending on the structural system. The code states that the damping value is a function of the frictional surfaces of the structural system utilized (e.g., welded steel structures might have lower damping values). ISO 4354 (ES ISO, 2009): This code recommends values for wind analysis for both steel and concrete structures for different/building heights and type of structure, as shown in Table 3.2. Chinese Standards (GB 50011, 2010): For seismic design, the code provides recommendations for steel structures: In case of frequent earthquakes, 4.0% for buildings with a height lower than 50 m; 3.0% for building height between 50 and 200 m; and 2.0% for building higher than 200 m. In case of rare earthquakes, the code suggests to utilize a 5.0% damping with an inelastic analysis. Hong Kong Code of Practice on Wind Effects (HKBD, 2004): The code recommends a damping ratio of 1.5% for the fundamental mode of steel structures and 2% for RC structures. If the building is particularly slender (i.e., high slenderness ratio), lower values than those mentioned could be suitable, in which case, the code recommends consulting with experts.

3.2 Damping in Tall Buildings

Table 3.2 Intrinsic Damping Ratio Recommended by ISO 4354 (ES ISO, 2009) Material Type of Structure Building

Chimney Lattice tower

H 5 40 m H 5 50 m H 5 60 m H 5 70 m H . 80 m

Steel (%)

Concrete (%)

1.8 1.5 1.5 1.5 1.0

2.0 2.0 1.5 1.5 1.2

0.2a 0.5b 0.1

1.0 

Note 1: The damping ratio is equal to the logarithmic decrement of damping divided by 2π. Note 2: Lower values (75% of the abovementioned values) are recommended for evaluation of habitability to horizontal vibrations of structures. a Unlined, all welded. b Lined.

•

US Standards • ASCE 7-16 (ASCE, 2017a): For seismic linear response linear analysis, a 5% maximum equivalent viscous damping ratio (for all modes with period greater than or equal to the period of vibration at which 90% of actual mass has been recovered in each orthogonal direction or response), unless proved by test data. In case of nonlinear analysis (with the inelastic behavior of elements modeled), the equivalent viscous damping shall not exceed 2.5%. Furthermore, the code states that damping shall be based on the material type, configuration, and behavior of the structure and nonstructural components, but it does not provide any recommended value or expressions to calculate it. For wind analysis, the code states that in the United States, 1%2% intrinsic damping is utilized for serviceability and 2.5%3% for ultimate strength design (based on several studies as described in the code). • ASCE 41-13 (ASCE, 2013): For seismic design, a 5%-damped spectrum shall be used, except in the following cases:  2% damping for buildings without exterior cladding  10% damping for buildings with wood diaphragms and a cross-wall with maximum of 40 feet in the direction transverse to the direction of motion  Effective damping according to the type of isolation or dissipation devices  Test data prove 5% cannot be used • PEER Tall Building Initiative (PEER, 2017): It recommends the use of equivalent viscous damping in the seismic response spectrum and seismic

69

70

CHAPTER 3 Damping considerations in tall buildings

linear/nonlinear history analyses if the inherent energy dissipation (e.g., due to structural elements or soilstructurefoundation interaction) is not explicitly modeled. The equivalent viscous damping can be assigned to higher modes with periods higher than 0.2 times that of fundamental mode, that is, Tm $ 0:2T1 , where m . 1. Alternatively, to ensure higher mode responses are not overdamped, mass- and stiffness-proportional Rayleigh damping model could be applied. The document provides recommendations for equivalent viscous damping of primary modes responses under both service-level earthquake (SLE) and maximum-considered earthquake (MCE), provided that no other values are justified based on evidence. For SLE, a limiting relation is given by (Fig. 3.18): 0:20 ζ cr 5 pffiffiffiffi # 0:05 H

(3.82)

where H is the building height in meters. For MCER, an equivalent viscous damping should not exceed ζ cr expressed in Eq. (3.82) and not less than 2.5%. Fig. 3.18 better illustrates the limit of damping for both SLE and MCER cases. A higher value of equivalent damping under MCER ground motion at longer building periods is recommended to take into consideration those energy dissipation effects

FIGURE 3.18 Equivalent modal viscous damping limit versus building height for SLE and MCER.

3.2 Damping in Tall Buildings

•

from soilfoundation interaction and elements undergoing inelastic regime. PEER/ATC 72-1 (PEER/ATC, 2010): An extensive description is given to the source of damping in tall buildings, and estimation techniques are given, similar to what described in previous sections. The document also states the importance of estimating the damping effects from soilstructure interaction, and they remark on the minimal information available on the subject. Usually deep foundation and rocking behavior provide higher damping, but it has also been found that tall buildings should have a low radiation damping component. In case of linear analyses, the guideline distinguishes between recommendations for wind and earthquakes: For wind, the following requirements are given: • Serviceability limit damping:  Steel-framed buildings: 0.5%1.0%  RC buildings: 1.0%1.5% • Strength limit damping:  Steel-framed buildings: 1.0%1.5%  RC buildings: 1.5%2.0% For earthquakes, recommendations are given for elastic analysis as 2% for steel-framed buildings (SAC, 1996) with a 5% upper limit based on the San Francisco Department of Building Inspection Administrative Bulletin, AB-83 (SFDBI, 2007). In case of nonlinear analyses, PEER/ATC (2010) suggests explicitly modeling the source of damping by utilizing Rayleigh or modal damping modeling. Moreover the following values of intrinsic damping are suggested: • Small deformations: 1% for steel frame structures and 2%3% for RC structures • Strong ground motions: 2%5% Alternatively a simple relationship for a building with a number of stories (N) is given for seismic analysis and design, as follows: α 30 α ζi 5 N

ζi 5

•

ðfor N , 30Þ ðfor N . 30Þ

(3.83)

where α is a coefficient that ranges between 60 (for steel systems) and 120 (for RC systems). These relationships show that it is recommended to have less damping for higher buildings. However, given the great uncertainty in the intrinsic damping estimation, it is advisable to conduct sensitivity studies. Document of Los Angeles Tall Buildings Structural Design Council (LATBSDC, 2017): The document states that equivalent viscous damping

71

72

CHAPTER 3 Damping considerations in tall buildings

•

•

•

•

ratio in tall buildings (at/below the yield level) can be considerably lesser than 5% that is often considered by building seismic codes. It recommends the same limiting values for equivalent damping as given by PEER (2017) (Fig. 3.18). This document encourages readers to refer to ATC-72 (PEER/ ATC, 2010) for more appropriate details about viscous damping for analysis. Austria O¨NORM B4014 Teil 1 (O¨NORM, 1993): It proposes that damping, for wind analysis, is composed of the summation of three contributions: • Structural damping ratio due to materials: 0.08% steel, 0.72% RC with cracks, and 0.40% RC without cracks • Structural damping ratio due to construction type: 0.32% steel tall buildings, 0.32% RC tall buildings (panel systems), and 0.64% RC tall buildings (frame systems) • Structural damping ratio due to foundations: 0.08% support with hinges, 0.24% support with sliding bearings, and 0.16% fixes support of frame structures British Standard BS ISO 4866:2010 (BS ISO, 2010): For wind analysis, Annex D states that no method for predicting intrinsic damping exists, and it is based on measured data in a range between 0.5% and 2.1%. However, higher values are encountered when significant soilstructure interaction is expected. Different methods have been proposed in ESDU 83009 (ESDU, 2012) and Jeary (1986), but their accuracy is not quoted. French Code (NV, 2009): In the wind section the code provides the intrinsic damping for the following categories (based on a logarithmic decrement): • 1.6% for steel • 3.2% for prestressed concrete structures • 4.8% for RC structures • 6.4% for masonry Italian Codes (NTC, 2008): In the NTC (2008) code, there are no specific requirements for wind design, but for earthquake, it implicitly states the use of 5% damping for all modes. Instead, in the CNR-DT 207-2008 (CNR, 2008), some formulations are provided for the inherent damping for the fundamental mode of buildings up to 200 m to be utilized for wind analysis, as follows: ζ i;1 5 ζ i;1 5

1 68 $ 0:01 ðfor RC buildings with h $ 30 mÞ 100 h

1 56 $ 0:008 ðfor steel buildings with h $ 30 mÞ 100 h

(3.84)

For higher modes (second and third), the following increase of inherent damping value can be considered: ζ i;m 5 1:4ζ i;1 ðfor i 5 2; 3 and RC buildings with h $ 50 mÞ ζ i;m 5 1:3ζ i;1 ðfor i 5 2; 3 and steel buildings with h $ 50 mÞ

(3.85)

3.3 Effect of Damping on Building

FIGURE 3.19 Fundamental intrinsic damping utilization for buildings in Japan. Adapted from Tamura, Y., Sasaki, A., Suda, K., 2000. Chapter 6: Design and damping evaluation, Damping in Buildings, Architectural Institute of Japan (in Japanese).

•

• •

Japanese Code (AIJ, 2000): The following values are recommended: • Steel buildings: 1% habitability (i.e., wind); 2% earthquake • RC buildings: 1% habitability (i.e., wind); 3% earthquake An interesting study was conducted by Tamura (2005) that shows the variation of damping utilized in the design of buildings in Japan (Fig. 3.19). Turkey Code (TDY, 2007; Dogangun, 2009): The earthquake code requires that a 5% modal intrinsic damping shall be considered for all modes. Canadian Building Code (NBCC, 2010): It states that for wind design, intrinsic damping should be based on experiments on real structures and that the common value for concrete frames is 2%.

It can be seen that the values proposed from the different codes vary significantly. Some of these values have been compared by Tamura (2005), and they are shown in Figs. 3.20 and 3.21.

3.3 EFFECT OF DAMPING ON BUILDING If supplementary damping is necessary for a structure, the effects of intrinsic and added damping both need to be considered. Compatible evaluation techniques, including those explained in Sections 3.4 and 3.5 and Chapter 5, should be considered in conjunction with other building performance objectives. Structural design performance evaluations are generally accomplished through a comparison of forces, movements, and accelerations. While more detailed numerical or laboratory simulations exist to predict the dynamic response of structures to wind effects, the following formulas offer a simple, normalized rule-of-thumb comparison to highlight the effects of added damping on buildings’

73

74

CHAPTER 3 Damping considerations in tall buildings

FIGURE 3.20 Code/guideline intrinsic damping comparison for RC buildings. Adapted from Tamura, Y., Suda, K., Sasaki, A., 2000. Damping in Buildings for Wind Resistant Design. Proc. International Symposium on Wind and Structures for the 21st Century, Cheju, Korea, pp. 115130.

potential responses before and after supplementary damping is considered. Across- and along-wind direction modifications are shown in the following: •

Across-wind direction

1=2 Rnew ζi 5 Rold ζ i 1ζ d

•

(3.86)

Along-wind direction "

1=2 # Rnew 1 ζi 12 512 2 Rold ζ i 1ζ d

(3.87)

where Rnew is the new response and Rold is the old response of a structure (e.g., force, deflection, or acceleration) which are function of the damping, with intrinsic (ζ i ) and supplementary damping (ζ d ) ratios. The above equations show how damping affects only the dynamic component of the motion. For along-wind direction motion, the static (mean) component is assumed to be half of the total (Eq. (3.87)). For across-wind direction, the motion does not have a static component (Eq. (3.86)).

3.3 Effect of Damping on Building

FIGURE 3.21 Code/guideline intrinsic damping comparison for steel buildings. Adapted from Tamura, Y., Suda, K., Sasaki, A., 2000. Damping in Buildings for Wind Resistant Design. Proc. International Symposium on Wind and Structures for the 21st Century, Cheju, Korea, pp. 115130.

Example #1: Consider an all-concrete shear core, 200 m tall residential tower, exhibiting an estimated 31 milli-g across-wind peak acceleration at a 10-year wind event, and 1.0% intrinsic damping (ζ i 5 1%) is assumed as part of the base building. If a 10-year acceptable acceleration limit is set at 18 milli-g, supplementary damping is estimated to be equal to 2% of critical (ζ d 5 2%) value to achieve this performance goal. The calculation is illustrated below, and the new acceleration response is 17.9 milli-g. 31

0:01 0:0110:02

1=2 5 17:9 milli-g

(3.88)

If a 20% reduction in along-wind overturning moment (OTM) direction was also targeted, from Eq. (3.87) an additional damping equal to 1.8% of critical (ζ d 5 1:8%) value would be required, as demonstrated in following: "

1=2 # OTMnew 1 0:01 5 0:80 5 1 2 12 OTMold 2 0:0110:018

(3.89)

75

76

CHAPTER 3 Damping considerations in tall buildings

Thus 3% total damping (1% 1 2%) would achieve the 18 milli-g acceleration performance goal, and 2.8% total damping (1% 1 1.8%) would achieve the performance goal of 20% overturning reduction. This would require both intrinsic and supplementary damping to meet these goals. The preceding is a rule-of-thumb example. More detailed techniques are currently available that refine such building behavior estimates and consider many variables that impact damping. Some of these frequently used techniques are described in a greater detail in Chapter 5. Another benefit of supplementary damping is the reduction in design sensitivity to large and inevitable uncertainty in a building’s intrinsic damping. Research by Tamura et al. (2000) suggests that there is a considerable side scatter in measured intrinsic structural damping. Furthermore intrinsic damping is thought to be dependent on the lateral system chosen, the height of the structure, the types of secondary systems and nonstructural partition wall assemblies, and the level of relative floor-to-floor displacement (see Section 3.2 for further details). The combination of these uncertainties increases the probability that the desired level of intrinsic damping might not be reliable. However, by providing a supplementary damping system, with an equivalent known percent of critical viscous damping, the uncertainty in total damping is reduced. Example #2: Assume an office tower with a concrete core lateral system and a steel-framed gravity system was designed based on an assumed intrinsic damping of 1.0%. Once the structure is built, intrinsic damping might reasonably vary between 0.5% and 1.3%. If the achieved damping was close to 0.5%, the actual wind acceleration could exceed the prediction by roughly 40%. If, however, the same building was designed assuming 1.0% intrinsic damping plus 1.5% supplementary damping, for a total of 2.5% damping, a 20.5% to 10.3% variation in intrinsic damping would result in total damping between 2.0% and 2.8%. If the expected acceleration performance was based on 2.5% damping, occupant comfort acceleration would only exceed the prediction by 11% for the lower bound limit of 2% total damping, significantly reducing the uncertainty level from previous 40%. Fig. 3.22 illustrates this reduction in design and performance uncertainty offered by the addition of engineered supplementary damping. The preceding example shows the importance of considering the expected variability and sensitivity of a building’s damping-dependent response, particularly if supplementary damping is provided for the reductions of ultimate limit state wind effects. A relatively small improvement in total damping can result in a significant reduction to the ultimate wind demands for a structure. For current tall building designs, supplementary damping is often employed to reduce human perception of motion and is rarely used for force reduction in the building’s design. It is debated in the industry whether supplementary damping should be used to reduce life-safety design demands. The designer needs to carefully consider the damping system’s long-term reliability. Indeed, buildings that include supplementary damping have the potential to have reduced damage and achieve better overall building performance on the condition that the dampers

3.3 Effect of Damping on Building

FIGURE 3.22 Sensitivity of occupant comfort performance to damping uncertainty.

would function as expected. If the damping is used to justify reduced structural quantities for life-safety calculations, it is critical that the damper system durability will not reduce overall structural reliability (see Sections 3.7 and Chapter 7).

3.3.1 OTHER DEVELOPMENT CONSIDERATIONS Damping systems, either distributed or discretely located, require planning, budgeting, appropriate procurement duration, and integration with remaining program spaces (including architectural, MEP, and other disciplines). Distributed systems integrated into the lateral system (see Section 4.1) can offer relatively compact placement, but it is possible that mechanical, electrical, and plumbing (MEP) systems may need to be rerouted to avoid the damper locations. Typically MEP systems cannot penetrate or interfere with the damping device or its anchorage. Cladding and secondary finishes in and around dampers may also need careful consideration, especially if concentrated movement locations that increase damper efficiency are included in the design of the building (such as an outrigger damper as explained in Chapter 4). Discrete systems, such as a roof-mounted pendulum dampers or sloshing tanks, are often easier to integrate into the building, but they require a relatively large volume of space near the tower top. They are also relatively heavy

77

78

CHAPTER 3 Damping considerations in tall buildings

(see Section 4.1.2.1), and often require placement near specific points of adequate load-carrying ability (e.g., above a reinforced elevator core). In either case, procurement, calibration, and maintenance costs of the damping system should be considered against the structural performance gains. Damping systems rely on relative movement between structural elements, or movement of the building itself. For this reason different damping devices such as viscoelastic or hydraulic axial dampers are often placed at locations with the largest differential movement, such as at outriggers. Absolute-displacement dampers such as tuned mass dampers should be placed at locations with high absolute movement, which are typically near a tower rooftop. Locating a damping system away from points of maximum movement can markedly decrease the damper’s efficiency. Absolute-displacement damper designs, which are sensitive to the building periods, are typically not finalized until a tower is built, and building measurements are taken to verify the actual building periods. Later system tuning may also be required should the building period change (e.g., after a seismic event where the building may soften). Contractors as well as selected subcontractors involved with damper systems often need to work to tighter tolerance requirements than other building elements and materials. Unlike typical concrete and steel tolerances defined by project specifications, a supplementary damping system is more like the installation of machine components where a higher level of construction precision may be required. For further details on these topics, see Chapters 6 and 7.

3.4 TALL BUILDINGS WIND-EXCITED MOTION The dynamic response of tall buildings can be influenced by many factors such as site conditions, shape, height, and slenderness of the building as well as dynamic characteristics of the structure (mass, stiffness, damping). The four main aspects in the design of a tall building that can be influenced by dynamic wind-induced excitation are: • • • •

Foundations and lateral-stability system Global deflection and inter-story drift Comfort of occupants at highest floors Performance of vertical transportation systems

Wind in buildings can induce two types of motion components: static or sustained action (utilized to estimate the building drift) and oscillatory or resonant vibration (due to the dynamics of wind loads). Both can lead to occupants’ discomfort when they become large. Motion is generally composed of sway in two horizontally perpendicular directions and torsion. Broadly speaking, there are two main types of mechanisms where wind can excite a tall building: buffeting and vortex shedding. The former—particularly

3.4 Tall Buildings Wind-Excited Motion

relevant in the along-wind/drag direction—is relatively benign and easy to predict following a codified approach. The latter—principally acting in a direction perpendicular to the one of the wind—has the potential to lead to a more substantial dynamic load amplification and being influenced by the architectural form of the building itself, could be much more accurately estimated through boundary layer wind tunnel testing. For tall or supertall buildings, the critical wind speed for vortex shedding is likely to be close to the 50- or 100-year return period events of the site, leading to potentially high dynamic ultimate limit state (ULS) design wind loads. In the case of much more flexible and slender structures, the resonant behavior is likely to occur at lower return period wind speeds and therefore affect the serviceability limit state (SLS) performance of the building. This is especially true for the new generation of super-slender buildings currently designed in several parts of the world, which often exhibit a resonant behavior with the vortex shedding excitation during events with return periods, sometimes as low as 1 month. There is no doubt that from an SLS standpoint, drift checks under wind loading are very important, but one should not forget that wind-induced acceleration checks can, in some cases, be more stringent. For example, if we consider a 200 m residential tall building with a natural frequency of 0.25 Hz exhibiting a total building drift of 0.35 m under 50-year return period winds, this would meet the widely used H/500 limit. If the corresponding total drift under 10-year return period winds is, say, 0.25 m and assuming that 50% of this is governed by dynamic behavior—which, in the across-wind direction, is not untypical—then the resulting wind-induced peak acceleration would move toward a rather undesirable 30 milli-g [50% 3 0.25 m 3 (2 3 π 3 0.25 Hz)2]. For particularly flexible and low-damped structures, it is therefore not uncommon for wind-induced vibrations to become one of the most challenging factors of the design. When it comes to the wind-induced structural response of tall buildings, damping is in fact one of the most influencing parameters controlling the level of dynamic load amplifications. For example, a residential tall building experiencing a level of wind-induced peak acceleration of 15 milli-g during a 10-year return period wind storm, at an estimated 2% total damping (fraction of critical) would meet the industry best practice criteria, but would fail to meet the performance goal if the level of damping of the structure only reaches 1%.

3.4.1 BUILDING WIND VIBRATION Dynamic wind-induced vibrations in tall buildings are the result of turbulent winds interacting with the structure at its natural frequencies. In particular vortex shedding resonant behavior typically occurs within rather specific wind speed ranges, and it is influenced by the architectural form of the tall building itself. In this phenomenon, alternating vortexes are shed on the downstream sides of tall buildings leading to net fluctuating forces exciting the structure in a direction that is perpendicular to the one of the wind (see Fig. 3.23 for a typical example of

79

80

CHAPTER 3 Damping considerations in tall buildings

FIGURE 3.23 Velocity contours along and around the faces of a tall building under wind oscillation [contours of relative velocity angle (degrees) (time = 1e-05)].

vortex shedding in a tall building analysis). When these forces occur at frequencies close to that of the structure, they may be significantly amplified, especially if the inherent damping of the structure is low. Compared to irregular buildings, monolithic prismatic buildings with uniform elevations are likely to exhibit stronger wind-induced oscillations due to vortex shedding. Unlike earthquake vibrations, wind-induced vibrations are dependent of the building’s three-dimensional form and/or cross-sectional uniformity or lack thereof. Therefore an effective strategy to mitigate wind oscillations in buildings involves the selection of the right plan-form configuration (see Section 3.8 for further details).

3.4.2 OCCUPANT COMFORT The parameter that is typically used to measure the performance of a structure undergoing wind-induced vibrations is acceleration because people do not sense building motion through displacement but they perceive it through the rate of change of the velocity involved with the motion itself. Wind-induced accelerations are conventionally reported as fractions of the acceleration of gravity (milli-g).

3.4 Tall Buildings Wind-Excited Motion

In order to help the readers of this document gain a better feeling of what certain levels of “milli-g” actually mean for a human being, the following extract from a monograph on wind-induced motion of tall buildings, published by the American Society of Civil Engineers (ASCE, 2015), is provided: •

• •

5 milli-g: perceptible to some occupants but, provided that such building motion does not occur frequently or continuously for an extended period of time, unlikely to cause significant adverse occupant response or alarm 10 milli-g: perceptible to the vast majority of occupants 3540 milli-g: a fear and safety threshold, sufficiently severe enough to cause some occupants to lose their balance

There is no doubt that wind-induced vibrations can be quite unnerving to the people occupying the uppermost levels of a tall building and could also trigger responses very similar to those associated with motion sickness, with symptoms ranging (depending on the severity of the motion) from concern, anxiety, fear, vertigo, dizziness, or headaches to nausea. Some of these symptoms are often enhanced by the presence of audio stimulus, such as creaking noises due to building sway and/or visual stimulus, which is particularly relevant in the case of a tall building exhibiting highly three-dimensional modes of vibration as a direct consequence of high eccentricities within the structural arrangement. The response of human beings to wind-induced building motion is a rather complex phenomenon involving many physiological and psychological factors. Perception (what people feel) and acceptability (what people are prepared to tolerate) are at the core of this important topic. The threshold of motion perception and its perceived severity can, in fact, vary widely from person to person. It also depends on expectation and experience. It should also be noted that perceptibility of wind-induced motion is inversely proportional to the square root of the product of mass, stiffness, and damping; this means that, in order to halve the perceptibility, the “mass 3 stiffness 3 damping” quantity needs to be increased by a factor of four.

3.4.2.1 Code perception criteria A significant amount of research work has been carried out over the past 40 years on the subject of human comfort in tall buildings subjected to wind excitation, particularly with the aim of determining suitable threshold levels for perception. Most of these studies have been carried out using motion simulators. Unfortunately in most of these experiments, unidirectional sinusoidal excitations were employed; hence, it is somewhat difficult to extrapolate the biaxial and torsional narrow-band random excitations experienced in an actual tall building (ASCE, 2015). When it comes to perception criteria, the design community in North America has traditionally focused on the structural performance during a typical 10-year return period wind storm: 1015 milli-g peak horizontal accelerations at the top occupied level for residential buildings and 2025 milli-g for offices have been

81

82

CHAPTER 3 Damping considerations in tall buildings

FIGURE 3.24 Comparison of average perception thresholds (where not indicated, the return period is not specified by the relative code).

widely used. The first codified criteria were introduced in the National Building Code of Canada (NBCC, 2010); again, focused on the performance during a 10-year return period wind event, the limits for peak horizontal accelerations were set at 10 and 30 milli-g for residential and office developments, respectively (Fig. 3.24 and Table 3.3). These limits are primarily adopted for frequencies between 0.15 and 0.3 Hz and are not essentially suitable for unusually high buildings or those with low frequency (ASCE, 2015). Unfortunately the above two criteria set threshold limits which are fixed, that is, they do not take into account the building frequency of oscillation. This is quite crucial as there is evidence that with decreasing frequency of oscillation, perception levels increase. Furthermore assessing occupant comfort for an event with a return period of 10 years cannot appropriately address the discomfort associated with far more regularly occurring wind storms. In general it is recommended that accelerations occurring with a 1-year recurrence interval should be reviewed because this shorter recurrence interval is far more relevant to the daily lives of the occupants. The Architectural Institute of Japan (AIJ), in 1991 (AIJ, 1991) (then revised in 2004 (AIJ, 2004)), proposed a new set of recommended levels of peak

3.4 Tall Buildings Wind-Excited Motion

Table 3.3 Comparison of Average Perception Thresholds Perception Threshold Standard

Return Period

Type

ISO 10137 (2007)

1-year

Residences

Office

AIJ (2004) CNR DT 207-2008 (CNR, 2008)

1-year

H90

H70

H50

H30

H10

NBCC (2010)

10-year

Residences Office

AS/NZS 1170.2 (AS/NZS, 2011) MGSN 4.19-05 (MGSN, 2005)

1-year 1-year

Not specified Not specified

Frequency (Hz)

Acceleration (milli-g)

0.06 1.0 2.0 5.0 0.06 1.0 2.0 5.0 0.1 1.5 2.5 5.0 0.1 1.5 2.5 5.0 0.1 1.5 2.5 5.0 0.1 1.5 2.5 5.0 0.1 1.5 2.5 5.0 0.15 0.30 0.15 0.30 Not specified Not specified

21.4 6.1 6.1 15.3 14.3 4.1 4.1 10.2 12.9 3.3 3.3 5.8 9.0 2.3 2.3 4.0 6.9 1.8 1.8 3.2 5.5 1.4 1.4 2.4 3.9 1.0 1.0 1.7 14.3 14.3 25.5 25.5 10 8

83

84

CHAPTER 3 Damping considerations in tall buildings

accelerations which, for the first time, were varying with the natural period of the tall building. These recommendations focused on 1-year return period wind events, which was found to be the most appropriate return period to evaluate motions in both tall and supertall structures (ASCE, 2015). The benefits of the AIJ recommendations are that they work around a framework which is performance based and set on five performance levels (Fig. 3.24 and Table 3.3; H-10, H-30, H-50, H-70, H-90, where the numbers represent the percentage levels of population that would perceive the motion). The selection of a particular criterion over the others is the result of a dialog between designers and the owner to define the desired habitability performance level. The same philosophy was adopted by the Italian norm CNR DT 207-2008 (CNR, 2008). Following the AIJ philosophy, in 2007, the International Organization for Standardization (ISO) published a set of guidelines, which have now become the most recognized criteria for habitability design internationally (ISO10137, 2007). The acceleration limits (again set against a 1-year return period) vary with the natural period of the structure, and they are more stringent for residential developments than for office buildings (Fig. 3.24 and Table 3.3). The Australia and New Zealand wind standard (AS/NZS, 2011) defines a simple criterion of 10 milli-g (Fig. 3.24 and Table 3.3) without specifying the wind speed and return period. According to Melbourne and Cheung (1988), this limit value refers to a 1-year return period. Similarly the Russian Standard (MGSN 4.19-05, 2005) sets an acceleration limit of 8 milli-g (Fig. 3.24 and Table 3.3) with a 1-year return period associated with tall buildings. As designers and owners often do not have a sense of what a specific level of “milli-g” actually feels like, in recent years motion simulators (ASCE, 2015), traditionally used by the maritime industry, have been employed to conduct experiments where the motion of the “moving room” has been driven by time histories obtained from the analysis of boundary layer wind tunnel testing. While this process is, of course, far from being robust and rigorous from a statistical viewpoint, it has helped designers, owners, and wind engineers through the decision-making process.

3.4.3 WIND DEFLECTION CRITERIA In addition to occupant comfort also deflection can poses several problem in a tall building due to wind excitation. Based on CTBUH (2008) deformations in a tall building are defined as: • •

Overall building movements: They are considered for the evaluation of the building overall performance such as P-delta effects. Story drifts: For the assessment of story performance, high-rise buildings have two types of displacement components: rigid body and panel (shear) deformations. The first one comes from the building rotation due to column/ wall vertical deformations. The second one is the in-plane deformation of a wall cladding panel (expressed with angle β in Fig. 3.25A, while θ is the story

3.5 Tall Buildings Earthquake-Excited Motion

FIGURE 3.25 Deformation in tall buildings.

•

drift ratio). In general for low-rise buildings these two deformations are equal while for high-rise buildings they differ, and their magnitude ratio depends on the structural system utilized (Figs. 3.25B and 3.25C). Inelastic deformations (particularly relevant for seismic considerations, Section 3.5): essential to assess the performance of ductile elements; for nonductile elements, none or little inelastic deformations are permissible.

Code limits on the lateral deflection of tall buildings are usually defined based on typical low- and medium-rise buildings, and their application to tall buildings poses several limitations. Moreover high-rise structural systems (outrigger systems, diagrid, bundle tubes) differ quite substantially from typical low-rise building structural systems, leading to different structural deflection performances. Furthermore, different performance deflection criteria should be defined for structural and nonstructural items. For these reasons performance-based design approaches are becoming common practice in the structural design of tall buildings (Smith, 2011). These methods define deflection limits based on the performance of several elements: structural systems, cladding and fac¸ade, interiors (e.g., walls and finishes), elevators, and occupant comfort. Most of the standards worldwide do not provide guidance on the overall deflection limits and are limited to inter-story drift (Table 3.4; Smith, 2011). It is a common practice to assume a limit of H/500 to H/300, where the normal return period is associated with the design wind loading as 50 years (not for climates with typhoons or hurricanes) or 100 years (for climates with typhoons or hurricanes) (Smith, 2011).

3.5 TALL BUILDINGS EARTHQUAKE-EXCITED MOTION During an earthquake, the performance of a building is mainly affected by the magnitude of its lateral inter-story displacements (drifts) and the acceleration of

85

86

CHAPTER 3 Damping considerations in tall buildings

Table 3.4 Code Wind Deflection Limits Return Period in Years

Building Type

JGJ3-2010 (JGJ, 2010)



JGJ 99-98 (JGJ, 1998) DG/TJ08-015 (2004)



Concrete/ Steel/ Composite H . 250 m Steel



Composite

Standard

Interstory Drift Ratio Limit

Overall Deflection Ratio Limit

1/500



1/400



1/500



Chinese

H . 250 m Hong Kong 



50 50

Concrete/ Steel Concrete Steel

 1/400

1/500 1/500











Steel

1/300





Concrete

1/500

10/50/100 (Engineering Judgment)

Steel

1/400-1/600





Steel

1/200



HKBD (2004)

50

HKBD (2013) HKBD (2011) Eurocode CEN (2010) British BS 5950-1 (1990) BS 8110-2 (1985) American ASCE 7-16 (ASCE, 2017a) Japanese AIJ-GBV (2004)

its floors. Drifts are related to both structural and nonstructural damage, while floor accelerations are related to damage to some of the structural and nonstructural components, as well as the contents of the structure. Traditional codes are usually not suitable for the seismic design of tall buildings since they are developed for low- and medium-rise buildings whose response is usually dominated by the first translational mode (in each direction). That is

3.5 Tall Buildings Earthquake-Excited Motion

not the case for tall buildings in which several translational modes contribute (in each direction) to the global response. For this reason, for tall buildings, a performance-based design (PBD) philosophy has been introduced in the last decades [starting in late 1990s with SEAOC (1995), ATC (1996), and FEMA (2000d)] to demonstrate satisfactory seismic performance. This method will allow the designer to assess how the building will perform during an earthquake of different levels, leading to more cost-efficient structural solutions (CTBUH, 2008; FEMA, 2012; PEER, 2017). In the following sections, there will be a brief review of PBD design philosophy, influence of damping in seismic building response, and deflection criteria to highlight the major aspects of seismic performance assessment of tall buildings.

3.5.1 PRINCIPLES OF PERFORMANCE-BASED EARTHQUAKERESISTANT DESIGN FOR TALL BUILDINGS PBD is a process for designing and retrofitting of existing buildings, to seek different performance goals for different earthquake intensities. Performance is usually expressed as probable (due to inherent uncertainties) damage and relative consequence from an earthquake, and the following performance measures can be utilized (FEMA, 2012): casualties, repair cost and time, and unsafe placarding. In addition, the possible performance goals to be achieved while designing a structure could be the following: • • •

Resist minor-intensity earthquakes without damage to structural and nonstructural elements or content Resist earthquakes of regular intensity without structural damage and limited nonstructural and content damage Prevent collapse during earthquakes of exceptionally severe intensity, even though they exhibit some structural, nonstructural, and content damage

Building codes do not explicitly verify these performance goals. Instead they provide “minimum” design requirements with seismic performance factors to provide safety against collapse in an earthquake (ASCE, 2017a; FEMA, 2009, 2012). For example, for ASCE 7-16 (ASCE, 2017a) the conditional probability of collapse ranges from 10% to 2.5%, depending on the building risk category (as reviewed in Chapter 5). The majority of structures, designed to meet modern building code requirements, have exhibited adequacy to meet the life-safety performance goal, as has been shown by recent earthquakes. However, major economic losses were caused mainly by nonstructural and building content damage rather than structural damage (FEMA, 2011). Therefore while considered a safe approach, code-based design has its disadvantages and limitations, including the following: •

Difficulties in predicting the expected structural and nonstructural damage levels for a particular ground motion intensity

87

88

CHAPTER 3 Damping considerations in tall buildings

• •

Unknown margin of safety against collapse Design provisions less suitable when extrapolated to tall buildings, because they are calibrated for low- to mid-rise buildings

Many of the limitations of the code-based design methodologies are caused by designing a structure that will respond in the nonlinear inelastic range, based on seismic performance factors that lead to achieve a life-safety performance objective with a low probability of collapse. PBD is a methodology that goes beyond these requirements. In a PBD approach, the performance goal assessment must be demonstrated in a rational and explicit manner, by showing that it satisfies the performance objectives expected of the structure. The current PBD approach originated in the 1990s from FEMA (1997) and has evolved into the current standards, namely, ASCE 7-16 (ASCE, 2017a) and ASCE 41-17 (ASCE, 2017b), and guidelines, namely, FEMA P-695 (FEMA, 2009), FEMA P-58 (FEMA, 2012)). The general methodology can be summarized in five tasks (FEMA, 2012): 1. 2. 3. 4. 5. 6.

Define performance goals Assemble building performance model Define earthquake hazard Analyze building response Develop collapse fragility function Calculate performance

Performance goals are defined as a series of discrete levels as operational, immediate occupancy, life safety, and collapse prevention. The process is initiated with a selection of one or more performance objectives by the design team, owners, and other stakeholders. Afterward, the building performance assessment is evaluated to predict the building response to earthquake hazards and to determine the probable consequence of damage. The design is then verified if the performance objectives are satisfied; otherwise it should be iterated until verification of performance goals. Based on FEMA (2009, 2012) performance assessment procedures can be divided into the following: •

•

•

Intensity-based: evaluate the building probable performance for a given earthquake shaking intensity; it is utilized for assessing the performance based on response spectrum. Scenario-based: evaluate the building probable performance for a given earthquake scenario based on a given building site location; this method is similar to the intensity-based method except that it takes into account the uncertainty in earthquake intensity. Time-based: evaluate the building probable performance over a specific period of time under the associated earthquake probability of occurrence; this method takes into consideration uncertainties in earthquake magnitude an location as well as the intensity of motion.

3.5 Tall Buildings Earthquake-Excited Motion

Additional details about PBD is beyond the scope of this publication and interested readers should refer to the relevant literature on the subject.

3.5.2 THE ROLE OF DAMPING IN THE SEISMIC RESPONSE CONTROL OF TALL BUILDINGS A straightforward approach to minimize lateral displacement in a building is to increase its lateral stiffness. However, stiffening a structure diminishes its performance on the side of content protection, since a more rigid building tends to have greater floor accelerations under seismic loading. As a result, there is a trade-off between accelerations and displacement demands, and improving one of these two parameters will result in diminishing the performance associated with the other one. Both of these aspects are shown schematically in Fig. 3.26 for elastic response spectral acceleration. A more effective strategy to reduce the response of a building is by increasing its total damping ratio. Unlike the traditional approach described before, this can reduce both displacements and accelerations at the same time. This is illustrated (for an elastic response behavior) in Fig. 3.27, where a reduction of both parameters for a given natural period results in a reduction of structural, nonstructural, and potential building content damage. Note that, Fig. 3.27 is valid only to only pseudo-response and not to maximum values (acceleration, velocity, and displacement). Moreover in case the structure is behaving inelastic, effective properties shall be utilized with an equivalent single DOF approach (Ramirez et al., 2001).

FIGURE 3.26 Effect of building stiffness on spectral acceleration.

89

90

CHAPTER 3 Damping considerations in tall buildings

FIGURE 3.27 Effect of damping on spectral acceleration.

3.5.3 SEISMIC DEFLECTION CRITERIA Building deformation is a crucial parameter to be estimated in PBD since it relates with the level of damages. As already explained for wind deflection criteria (in Section 3.4.3), there are three deformation parameters: overall building movements, story drifts, and inelastic deformations. The same principles already explained are applied to seismic deflection, and the only major difference is related to the inelastic deformations. As explained in the previous sections, the building may experience some inelastic deformations during strong events without leading to any collapse. Therefore code limits on the lateral deflection of tall buildings under seismic loads are higher than those induced from wind loads (Section 3.4.3). However, similar to wind design, most of the standards worldwide do not provide guidance on the overall deflection limits and are limited to the inter-story drift ratio for linear analyses (Table 3.5; Smith, 2011).

3.6 ENVIRONMENTAL AND ECONOMIC CONSIDERATIONS The economic considerations involved in the utilization of dynamic modification devices are a function of the performance objective that the structure needs to be designed for (as explained in the previous sections). Moreover in the case of extremely hazardous events (e.g., hurricanes, earthquakes, etc.), the level of damage accepted for the building has an influence on the level of costbenefit for each solution. In this way each design option can be compared in terms of probability of exceeding loss versus overall cost. To perform all these evaluations, consideration of the environmental aspects of the building life cycle is also

Table 3.5 Code Seismic Deflection Limits

Standard

Return Period in Years

Interstory Drift Ratio Limit

Overall Deflection Ratio Limit

Concrete/steel/ composite H . 250 m Steel

1/500



1/250



Composite H . 250 m Multistory and tall steel structure Reinforced concrete frame Reinforced concrete seismic wall, tube-in-tube/ supported story Reinforced concrete frame-seismic wall, and framecore-tube Reinforced concrete seismic wall, tube-in-tube/ multistory and tall steel structure Reinforced concrete seismic wall, tube-in-tube Reinforced concrete frame-seismic wall, and framecore-tube

1/500



Building Type

Chinese r3-2010 (JGJ, 2010)

51a

JGJ 99-98 (JGJ, 1998) DG/TJ08-015 (2004)

51a

GB50011 (2010)

51a 51a 51a

51a

51a 16422475b 16422475b 16422475b

1/250 1/550 1/1000 1/800 1/50



1/120 1/100

 

1/200



Eurocode Eurocode 8 CEN (2003)

475c

475c 475c

Steel/concrete Buildings including nonstructural members with brittle material connected to structure Steel/concrete Buildings including ductile nonstructural members Steel/concrete Buildings without nonstructural members or including fixed nonstructural members

1/133 1/100

(Continued)

Table 3.5 Code Seismic Deflection Limits Continued

Standard

Return Period in Years

Building Type

Interstory Drift Ratio Limit

Overall Deflection Ratio Limit

1/50d



American ASCE 7-16 (ASCE, 2017a)

475c

ASCE 41-17 (ASCE, 2017b) FEMA P-1050-1 (NEHRP, 2015)

475c

FEMA P-1050-1 (NEHRP, 2015) PEER (2017), LATBSDC (2017)

Steel/concrete Risk category I or IIe Steel/concrete

2475f

Risk category IIIe Steel/concrete Risk category IVe Steel/concrete

43g 2475f

Steel/concrete Steel/concrete

100 500 100 500

Steel

c

475

1/67d

1/100d 1.5 times the limits given by ASCE (2010) 1/200 1/33h 1/22i 1/100j 1/67k

Japanese AIJ-GBV (2004)

a

Concrete

Frequent earthquake (63% exceedance probability in 50 years). Rare earthquake (23% exceedance probability in 50 years). Design-level earthquake (10% exceedance probability in 50 years). d In the case of nonlinear analysis, the limits are doubled. e See more details in Section 5.2.1.1.1. f Maximum-considered earthquake. g Service-level earthquake (50% probability of exceedance in 30 years). h Mean of the absolute values of the peak transient drift ratios. i Absolute value of the maximum story drift ratio from the suite of analyses. j Mean of the absolute values of residual drift ratios from the suite of analyses. k Maximum residual story drift ratio in any analysis. b c

1/200 1/75-1/100 1/500 1/200



3.6 Environmental and Economic Considerations

important. The best tool to serve this purpose is life-cycle assessment (LCA). This methodology will allow the owner to choose among the different options based on costs, sustainability, and the level of performance that is acceptable. In the following section, the most relevant LCA theory and tools available in literature are briefly reviewed, together with some building examples. Interested readers should review the relevant literature on the subject for further details.

3.6.1 LIFE-CYCLE ASSESSMENT The construction industry has a significant role in global energy consumption. Several publications (UNEP, 2007; Pulselli et al., 2007; Na¨sse´n et al., 2007) have reported that the building sector contributes to as much as 40% of the world’s material consumption and 30%40% of the total energy demand and greenhouse gas emissions. This is why sustainable development is necessary for the construction industry, and this could be related to a triple bottom-line strategy (Willard, 2002). See such a strategy illustrated in Fig. 3.28. Developed in the 1990s, LCA is, to date, the most complete tool for assessing the building’s environmental impact (Fava et al., 1991; ISO 14040, 1997). LCA considers the building’s initial construction, usage, and end-of-life phases. Additionally it is necessary to include the environmental impacts that can occur

FIGURE 3.28 Triple bottom-line representation and sustainable requirements. Adapted from Menna, C., Asprone D., Jalayer, F., Prota, A., and Manfredi, G., 2013. Assessment of ecological sustainability of a building subjected to potential seismic events during its lifetime. Int. J. Life Cycle Assess., 18(2), 504515.

93

94

CHAPTER 3 Damping considerations in tall buildings

for repairing damage from potential natural disasters (e.g., earthquakes) (Menna et al., 2013). One of the major limitations of LCA is that it does not consider how impacts change over time (Trusty, 2008). LCA was standardized for the first time in 2006 (ISO 14040, 2006 and ISO 14044, 2006), and it consists of four phases (ISO 14040, 2006): •

•

•

•

The goal and scope shall be chosen to provide the most reliable and real results. The goal reflects the selection of the building’s life-cycle stages that shall be clearly motivated. Different components shall be considered in the LCA: the functional unit, the functions of the system, data requirements, impact categories, methodology of the impact assessment, and interpretation. The scope defines what are and how to utilize the functional units and how to interpret them. It is relevant in this phase to understand the data quality that is utilized (i.e., age, geographic location, variability, etc.). Inventory analysis consists of creating a work-flow diagram of the entire building life cycle. This is done by collecting all the data for environmental calculations, both input (e.g., energy, water, and land use) and output (e.g., emissions, use of nonrenewable resources) during the lifetime of the building. Impact assessments classify, grade, and group the different environmental lifecycle impacts. This phase is really important, in order to make the different factors uniform and comparable. One of the most utilized factors is carbon emissions expressed as CO2 equivalents. Interpretation of results allows the identification of major factors that impact the building’s sustainability and provides recommendations.

The building material/components for LCA purposes can be divided into the following categories (Menna et al., 2013): • • • • •

Foundation Superstructure Nonstructural members: siding, insulation, interior, trim, and door/windows Water and electrical systems Major appliances

The diverse building components have a contribution in the building’s initial cost. For example, studies by Taghavi and Miranda (2003) state that the structural system is 10%20% of the total construction costs. Therefore the majority of the costs are related to nonstructural components, and their values vary depending on the building function (Fig. 3.29). Moreover the mechanical system contributes to 20%30% of the construction cost while the electrical system only contributes to 10%. Despite this low impact on the total initial cost, structural components contribute significantly to the total energy consumed by the building, due to the high energy consumption during manufacturing and transportation of structural materials (Menna et al., 2013). The environmental considerations for each LCA category can be determined by looking at the energy consumption that can be divided into two

3.6 Environmental and Economic Considerations

FIGURE 3.29 Cost distribution by building type. Data from Taghavi, S., Miranda, E., 2003. Response Assessment of Nonstructural Elements. PEER Report 2003/05, Pacific Earthquake Engineering Research Center, Berkeley.

categories: embodied and operational. The former is the energy used to construct, transport, maintain, and dispose of the building. The latter is the energy utilized to keep the building functional (e.g., heating, water, lighting, etc.). The energy utilized during the occupation phase has been shown to be the major contributor (between 50% and 70%) (Peuportier, 2001; Scheuer et al., 2003; Athena, 2013). This is a function of the building life span, usually set as 66 years (Athena, 2013). The life span could be extended, especially in seismic areas, thanks to PBD (see previous sections). Typical embodied energy values are shown in Table 3.6 for different building materials. From the table, it can be seen that, among the most popular structural materials, steel has the highest embodied energy per ton, while concrete has the highest rate of CO2 emission (Webster, 2001). Fig. 3.30 shows the contribution of different factors to the total embodied energy of a typical office building. There are two major LCA approaches utilized in the construction industry (Ortiz et al., 2009; Erlandsson and Borg, 2003): bottom-up LCA for the combinations of material and components and top-down LCA of the whole construction process. LCA will also help in the decision-making phase of building construction. Low-impact buildings usually imply advance technological and innovative material and construction techniques that in most cases require additional upfront financial investments. In contrast, this could lead to savings for the building’s usage in extreme events (e.g., earthquakes, hurricanes), as will be reviewed in the following section.

95

96

CHAPTER 3 Damping considerations in tall buildings

Table 3.6 Embodied Energy for Typical Materials Material

MJ/kg (Embodied Energy per Unit of Mass)

Aggregate Straw bale Concrete block Precast concrete Concrete Lumber Brick Plywood Steel (new) Steel (recycled) Cellulose insulation Gypsum wallboard Linoleum Carpet Aluminum

0.10 0.24 0.94 2.00 1.30 2.50 2.50 10.40 32.00 8.90 3.30 6.10 116.00 148.00 227.00

From Kneer, E., Maclise, L., 2008. Consideration of building performance in sustainable design: a structural engineer's role. In: Proceedings of SEAOC Convention, pp. 115.

FIGURE 3.30 Embodied energy breakdown for a typical office building. Data from Cole, R.J., and Kernan, P.C., 1996. Life-cycle energy use in office buildings. Build. Environ., 31(4), 307317.

3.6 Environmental and Economic Considerations

3.6.1.1 Damaged-oriented life-cycle assessment Given the sustainability aspects of the building construction, critical actions coming from natural hazards (e.g., earthquakes, hurricanes) should also be considered. For this aim, building LCA that includes seismic damage was extensively studied in the last decade (Comber et al., 2012; Kneer and Maclise, 2008; Menna et al., 2013; Padgett and Tapia, 2013). The goal is to understand the importance of natural hazard mitigation for a building versus the environmental impact it would have. In particular Menna et al. (2013) proposed an LCA analysis approach that includes seismic damage based on their probability of exceedance. This study shows how seismic events can influence LCA from 6% to 25%, depending on if the whole building’s life or just the initial construction phase is considered. Therefore given the significant damage costs, reaching a higher performance level could be beneficial for the overall life-cycle costs of a building. A comparative study was conducted by Comber et al. (2012) between a concrete moment frame and a shear wall system. Results showed that, while the shear wall has more associated carbon outputs, when looking at the total embodied carbon, the difference between the two systems becomes substantial when considering the carbon output from seismic repair. Therefore the experts suggest the necessity of an environmental rating system (e.g., LEED) to promote disaster-resilient design strategies. In order to evaluate the environmental performance of damage on a structure, one possible approach is IMPACT 20021 , proposed by Jolliet et al. (2003), as shown in Fig. 3.31. This method links the midpoint categories to four different

FIGURE 3.31 Scheme of the IMPACT 20021 framework (LCI: life-cycle inventory). Adapted from Jolliet, O., Margni, M., Charles, R., Humbert, S., Payet, J., Rebitzer, G., et al., 2003. IMPACT 2002 1 : a new life cycle impact assessment methodology. Int. J. Life Cycle Assess. 8(6), 324330.

97

98

CHAPTER 3 Damping considerations in tall buildings

damage categories. This allows the identification of the possible environmental impacts on a building. Different types of damage can occur in buildings as a consequence of seismic events: structural, nonstructural (architectural, mechanical, and electrical), and contents. Architectural damage is correlated with inter-story drift and MEP (mechanical, electric, and plumbing systems) damages are correlated with both inter-story drifts and floor accelerations (FEMA, 2012). Several authors have tried to quantify and correlate damage to cost and environmental impacts. One of the studies was conducted by Ferritto (1984), which proposed an estimation of damage costs, based on two components (Tables 3.7 and 3.8): • •

Drift will impose damage to the main structure, as well as nonstructural components such as cladding, windows, partitions, etc. Floor accelerations will impose damage to ceiling and building contents.

When it is necessary to quantify the loss estimation due to earthquake-induced damages, it is important to differentiate between direct and indirect losses. The former relates to the cost associated with repair of building components, while the latter related to temporary loss of building function (downtime). Several methods can be utilized to estimate losses, such as: •

Regional loss estimation methods that utilize average building response from mean building characteristics and related ground motion intensity; these are utilized to analyze a large building stock and while simple might not be suitable for atypical buildings. One of the most utilized methods in the USA is the HAZUS Methodology and Software (FEMA, 2015). In this method structural demand parameters estimation is supported by fragility function. Fragility functions (Fig. 3.32) give the probability of exceeding a certain level of damage (i.e., slight, moderate, extensive, and complete damage) as a function of an engineering demand parameter such as inter-story drift ratio and floor spectral acceleration. As an alternative to the HAZUS approach, another method commonly utilized relies on the utilization of damage indexes that were mainly introduced for determining whether a building needs to be retrofit or not, without any quantitative evaluation of economic losses. One of the most utilized damage indices was proposed by Park and Ang (1985), as: DI 5

ð δmax β 1 dEH δu δ u fy

(3.90)

where δmax is the maximum deformation, δu is the ultimate deformation of the element, fy is the element yield strength, EH is the hysteretic energy of the element, and β is the model constant parameter. This index is calculated for each element and, through a weighting factor, the story and the total building damage can be calculated. Park et al. (1987) defined different limit states as shown in Table 3.9.

Table 3.7 Damage Ratios Due to Drifts for Steel Structures Story Drift (%)

Rigid frame Braced frame Shear wall Nonseismic frame Masonry Windows and frames Partitions, architectural elements Floor Foundation Equipment and plumbing Contents

Repair Multiplier

0.1

0.5

1.0

2.0

3.0

4.0

7.0

10.0

14.0

2.0 2.0 2.0 1.5 2.0 1.5 1.25 1.5 1.5 1.25 1.0

0 0 0 0 0 0 0 0 0 0 0

0.01 0.03 0.05 0.005 0.10 0.30 0.10 0.01 0.01 0.02 0.02

0.02 0.14 0.30 0.01 0.20 0.80 0.30 0.04 0.04 0.07 0.07

0.05 0.22 0.30 0.02 0.50 1.00 1.00 0.12 0.10 0.15 0.15

0.10 0.40 0.60 0.10 1.00

0.20 0.85 0.85 0.30 1.00

0.35 1.00 1.00 1.00

0.50

1.00

0.20 0.25 0.35 0.35

0.35 0.30 0.45 0.45

0.80 0.50 0.80 0.80

1.00 1.00 1.00 1.00

From Ferritto, J.M., 1984. Economics of seismic design for new buildings. J. Struct. Eng., 110, No. 12.

100

CHAPTER 3 Damping considerations in tall buildings

Table 3.8 Damage Ratios Due to Floor Acceleration for Steel Structures Floor Acceleration (g)

Floor and roof system Ceilings and lights Building equipment and plumbing Elevators Foundations (slab on grade, sitework) Contents

Repair Multiplier

0.08

0.18

0.5

1.2

1.4

1.5 1.25 1.25 1.5 1.5 1.05

0.01 0.01 0.01 0.01 0.01 0.05

0.02 0.10 0.10 0.10 0.02 0.20

0.10 0.60 0.45 0.50 0.10 0.60

0.5 0.95 0.60 0.70 0.50 0.90

1.00 1.00 1.00 1.00 1.00 1.00

From Ferritto, J.M., 1984. Economics of seismic design for new buildings. J. Struct. Eng., 110, No. 12.

FIGURE 3.32 Example of HAZUS-MH fragility curve. Adapted from FEMA, 2005. FEMA 400: NEHRP Improvement of Nonlinear Static Seismic Analysis Procedures. Washington, DC: Federal Emergency Management Agency  ASCE.

Table 3.9 Interpretation of Overall Damage Index Limit State Damage Index 0.00 0.200.30 0.500.60

.1.00

Degree of Damage

Damage (Service) State

None Slight Minor Moderate Severe Collapse

Undamaged Serviceable Repairable Unrepairable Collapse

From Park, Y.J., Ang, A.H.-S., and Wen, Y.K., 1987. Damage-limiting aseismic design of buildings. Earthquake Spectra, 3:1,12.

3.6 Environmental and Economic Considerations

•

Component-based loss estimation methods that estimate losses based on earthquake damage to each component in a building require a great amount of data and computational resources. In most of the cases, these methods require explicitly structural modeling to estimate the structural response parameters that can be then related to component damage. One of the most utilized approaches is the probabilistic evaluation of the seismic structural response (FEMA, 2012; Miranda and Aslani, 2003) based on the pioneering work by Cornell (1968) on Probabilistic Seismic Hazard Analysis (PSHA). This procedure aims to compute the annual probability of exceedance of various types of Engineering Demand Parameter (EDP) (e.g., peak inter-story drifts and floor accelerations). Then loss estimation can be carried out from the method originally developed by Cornell and Krwinkler (2000), which can be expressed by the following formulation (Aslani, 2005; Ramirez and Miranda, 2009): E½LT jIM  5 E½LT jNC; IM PðNCjIM Þ 1 E½LT jCPðCjIM Þ

(3.91)

where E½LT jNC; IM  is the expected loss of no collapse for a building under ground motion intensity IM, E½LT jC is the expected loss of collapse, PðNCjIM Þ is the probability of no collapse for a building for ground motion intensity IM, and PðCjIM Þ is the probability of collapse for ground motion intensity IM (which is complementary to PðNCjIM Þ 5 1 2 PðCjIM Þ). In addition to the economic loss, it is valuable to evaluate the environmental impacts of the seismic damage and, for this scope, a correlation between a building’s embodied energy-to-cost can be utilized (Feese, 2013). This can be obtained from dividing the building’s embodied energy by its cost and multiplying the damage costs considered during the life span of the building. Feese (2013) has shown from comparing low-rise steel- and concrete-frame buildings, that the seismic damage impact is small compared to the total embodied energy of the building (,1%). It is important to note that this study did not include nonstructural building items, which are, in most cases, major sources of economic loss. Further details about this topic are outside the scope of this book, and interested readers should refer to the relevant literature on the subject.

3.6.2 COST IMPLICATION OF BUILDINGS WITH DISSIPATION DEVICES The cost of supplementary damping devices can be divided into four categories (Tse et al., 2012): • • • •

Design and feasibility study Procurement of mechanical components and manufacturing On-site installation and commissioning of the system Maintenance of the system over a period of time

101

102

CHAPTER 3 Damping considerations in tall buildings

The cost of each device will vary depending on the capacity. For the maintenance costs (see Section 7.6.6), active and semiactive dampers will typically require higher costs than passive ones (Shin, 2010). However, higher costs are also associated with yielding metallic devices that may require replacement after a major event. Given the small contribution of the structural system, in terms of the total construction costs (10%20%, as shown in the previous section), it is clear that the additional initial costs due to dissipation devices are generally minor (Kneer and Maclise, 2008). Note that the cost comparison among different building schemes may also depend on the relative building material cost in different nations. The cost system in Asia will be quite different from that in North America.

3.7 DAMPING TECHNOLOGY UNCERTAINTY AND ROBUSTNESS PERFORMANCE The design of damping devices involves great variability in their parameters due to the uncertainties in the variables utilized in their design (e.g., building frequency, material properties, aging effects, etc.). Thus the optimization of damper parameters considering the model variability is an important issue in order to have a robust and reliable performance. Systematic procedures with bounding analysis (i.e., evaluation of parameters with proper lower- and upper-bound properties) for uncertain structural parameters have been successfully utilized in the past (e.g., Alefeld and Herzberger, 1983; Chen and Wu, 2004; Chen et al., 2009; Moore, 1966; Qiu, 2003) and are discussed in depth in Section 5.1.3.3.3. This is a common approach to reduce uncertainty and increase the building’s sustainability. However, given the high number of parameters, the combination can increase exponentially. This can be overcome by sensitivity analysis or Taylor series (e.g., Fujita and Takewaki, 2011a,b, 2012; Kanno and Takewaki, 2006; Takewaki, 2008a,b; Takewaki and Fujita, 2014). Furthermore it is important to find the critical combination of the interval parameters (Takewaki, 2015). Uncertainties also come from the estimation and prediction of hazard events. In seismic design, probability theories are often used to characterize their rate of occurrence and, currently, design approaches are moving on from the concept of “critical excitation” to include the ground motion variability (Takewaki, 2015). This would define the specific excitation parameters, depending on each building characteristics and the site under study. All this would induce a more reliable and robust seismic-resistant design. Dealing with uncertainties involves a lot of variables; therefore, it is important to define the type of strategy to utilize. Takewaki (2015) describes how there could be three major approaches: • •

Worst-case analysis: This analysis aims to find the maximum structural response of the system based on a class of variable inputs. Structural control and health monitoring: These methods permit reduction in the uncertainty of the structural parameters since the response is predicted

3.8 Alternative to Damping Devices

•

accurately through time (i.e., variability in time can then be studied). This could result in an elongation of the service life of buildings. Resilience: It is related with earthquake resilience and refers to structures that have reduced probability of collapse, damage, and shorter recovery time (Bruneau and Reinhorn, 2006). Bruneau and Reinhorn (2006) proposed four resilience measures: • Robustness: Devices properties do not suffer degradation or loss of function for any demand level • Redundancy: Elements capable of maintaining their functionality in the event of disruption, degradation, or loss of functionality • Resourcefulness: Recovery by utilizing the most suitable material and resources to achieve the goals • Rapidity: In order to contain losses, the recovery as fast as possible

Given the high level of uncertainties involved in the design of structures with damping technology, designers should be aware of these different approaches. Further considerations specific to evaluation of damper properties estimation for bounding analysis can be found in Chapter 5.

3.8 ALTERNATIVE TO DAMPING DEVICES The trend toward increasingly slender buildings can result in structures that are vibration sensitive and, therefore, more likely to require some form of mitigation to limit motion perception by the occupants. There are, however, several options available to building designers to achieve this. One approach, which is the focus of this book, is the addition of supplementary damping. Other approaches exist that have some practical limitations such as increasing the mass and/or stiffness of the structure beyond what is necessary for strength design and drift control. One technique that can offer significant benefits without the associated costs of supplementary damping devices, increased mass, or structural stiffness is the aerodynamic design of the building form. Substantial improvements to a building’s aerodynamic performance are possible through relatively small modifications to its form. The dynamic oscillations experienced by high-rise buildings are primarily a result of vortex shedding, occurring as air flows around them (as explained in Section 3.4). Modifications to the building form can reduce the sensitivity of the building to vortex shedding and, hence, the motions perceived by occupants. These modifications can be divided into two groups (Mooneghi and Kargarmoakhar, 2016): •

Minor modifications that have a minimal impact on the architectural and structural design of the building: simple modifications are corner interventions (Fig. 3.33) or changing the building orientation with respect to the primary wind direction. The effect of corner modification is a function of the size and type of corner utilized as well as the plan dimensions. Several studies have been conducted on this subject in the past (e.g., Gu and Quan, 2004; Irwin,

103

104

CHAPTER 3 Damping considerations in tall buildings

FIGURE 3.33 Minor aerodynamic modifications that improve the performance of the building. Adapted from Mooneghi, M.A., Kargarmoakhar, R., 2016. Aerodynamic mitigation and shape optimization of buildings: review. J. Build. Eng. 5, 225235.

FIGURE 3.34 Major aerodynamic modifications. Adapted from Mooneghi, M.A., Kargarmoakhar, R., 2016. Aerodynamic mitigation and shape optimization of buildings: review. J. Build. Eng. 5, 225235.

•

2008; Irwin et al., 2008), but each case study should be carefully evaluated to avoid unfavorable effects on the building behavior. Major modifications that significantly affect the structural and architectural design: the most utilized of these modifications are shown in Fig. 3.34.

3.8 Alternative to Damping Devices

FIGURE 3.35 Cross-wind overturning moment spectra for tall building models with various configurations. Adapted from Tanaka, H., Tamura, Y., Ohtake, K., Nakai, M., Hayano, Y., Koshika, N., 2009. Study on wind force characteristics for super high-rise building with new shape, (Part 1) Comparison of wind force characteristics of various building shapes. In: Summaries of Technical Papers, Annual Meeting, Architectural Institute of Japan, B-1, Structures I, (to be published) (in Japanese).

Studies (Tanaka et al., 2013, 2009) have been conducted on the wind-induced response of tall buildings with various configurations (Fig. 3.35). The most efficient surface shape could be determined through optimization techniques with wind tunnel and computational fluid dynamics (CFD) analyses (Mooneghi and Kargarmoakhar, 2016). However, a systematic approach does not yet exist for defining the global shape of tall buildings (Mooneghi and Kargarmoakhar, 2016). As this technique involves modifying the building form, it requires a collaborative design approach involving the architect, structural engineer, and wind engineer, early in the design process. It also requires a more extensive wind tunnel program, typically including a wind tunnel workshop, where the design team works collaboratively to test a multitude of potential

105

106

CHAPTER 3 Damping considerations in tall buildings

FIGURE 3.36 Burj Khalifa scheme comparison.

aerodynamic design enhancements before choosing an optimized aerodynamic solution. An equally important benefit to this approach is by focusing on reducing the dynamic wind loads that building experiences, rather than just mitigating vibrations; cost savings can also be realized through a reduction in structural materials required for the strength design of the structure. This aerodynamic design approach was employed in the design of currently the world’s tallest building, the Burj Khalifa (Fig. 3.36), to effectively reduce the design wind loads as well as eliminate the need for any supplementary damping devices. Potential future developments to this approach could be the active control of the aerodynamic performance through the use of deployable spoilers or shape-adjusting fac¸ades. In addition to currently applicable techniques described earlier, where the physical geometry of the tower is modified to enhance aerodynamic performance, a topic of recent research is aerodynamic enhancement through the active control of the boundary layer of fluid flowing around a solid object. This technique of Fluid-based Aerodynamic Modification (FAM) was developed for the aerospace industry, and recently, its application to building design was studied (e.g., Menicovich et al., 2012). This proposed approach would rely on a system of sensors and actively controlled actuators at the building fac¸ade, which serve to disturb and modify the flow of fluid around the building in such a way that the airflow virtually “sees” a different tower shape.

CHAPTER

An introduction to dynamic modification devices

4

CHAPTER OUTLINE 4.1 Passive Damping Systems ................................................................................108 4.1.1 Distributed Damping Approaches .....................................................109 4.1.2 Mass Damping Approaches..............................................................195 4.2 Seismic Isolation .............................................................................................207 4.2.1 Base-Isolation Types .......................................................................210 4.2.2 Base-Isolation Manufactures ...........................................................215 4.3 Active, Semiactive, and Hybrid Systems ............................................................217 4.3.1 Active Systems...............................................................................222 4.3.2 Hybrid Systems ..............................................................................224 4.3.3 Semiactive Systems........................................................................226 4.3.4 Adaptive Tuned-Mass Damper Systems.............................................229 4.3.5 Control Strategies ...........................................................................230 4.3.6 Future Directions............................................................................231 4.3.7 Active, Semiactive, and Hybrid Dampers Manufactures......................232 4.4 Comparison of Dampers in Tall Buildings...........................................................232

Dynamic modification device technology has its origin in tall buildings since the early 1970s, but has only started to be more widely implemented in the past few decades, as explained in Chapter 2. Choosing which system to utilize depends on many variables, including, but not limited to, the specific structural system, building heights, and external excitation. It is important to understand that there are benefits and deficiencies of each system, which determine the criteria for their selection and design. In this section, the basic mechanical and design properties of each system, which are independent of the structural system, are reviewed, and the type of building they will be generally utilized on is identified. Among the different solutions available for supplemental damping and seismic isolation systems, there are three main categories: passive, active/semiactive/ hybrid, and seismic isolation systems. Passive systems differ from active/semiactive/hybrid systems, as they do not need any external power source to be activated. In order to be activated, passive dampers typically rely on the natural movement of the structural system. The seismic isolation systems are considered independent from the first two, since the main function of the system is to decouple the building response of the structure above the isolation level. To achieve Damping Technologies for Tall Buildings. DOI: https://doi.org/10.1016/B978-0-12-815963-7.00004-X Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved.

107

108

CHAPTER 4 An introduction to dynamic modification devices

this, the lateral stiffness of the superstructure (because of the isolation system) is lowered, so that the period is longer. This moves the building farther out of resonance with the earthquake ground motions, so that the energy transferred from the ground motion to the superstructure can be reduced. If isolators do not provide enough additional damping, it can be coupled with supplemental damping devices. These three main categories have different devices, summarized in Table 4.1. In this chapter, these categories are reviewed in detail identifying their mechanical properties and basic design principles.

4.1 PASSIVE DAMPING SYSTEMS Passive energy-dissipation systems are generally divided into three different categories: displacement-dependent, velocity-dependent, and motion-dependent devices (Table 4.2). However, there are special devices, not belonging to any of these categories: viscoelastic (VE) dampers, which are in between displacementand velocity-dependent devices, and friction dampers, which are activated at a certain force value that does not depend on velocity or displacement. Displacement-dependent dampers do not depend on the frequency of motion, and the forces generated by these devices are usually in-phase with the inertial Table 4.1 Typologies of Damper Devices Used in Practice Passive Dampers Metallic Friction Viscoelastic Viscous Tuned mass Tuned liquid Tuned column liquid Self-centering

Semiactive/Active/ Hybrid Dampers Diagonal braces Tuned mass Variable stiffness Variable damping Piezoelectric Rheological

Isolation Systems Low-damping rubber High-damping rubber Lead rubber Metallic Lead extrusion Friction pendulum Rotating ball bearing Helical steel springs plus VE dampers

Table 4.2 Categories of Passive-Added Damping Devices DisplacementDependent

VelocityDependent

MotionDependent

Metallic Self-centering

Viscous

Tuned mass Tuned liquid Tuned column liquid

Others Friction Viscoelastic

4.1 Passive Damping Systems

forces from the external excitation. The energy dissipated is a function of the relative displacement between the damper ends. In contrast, velocity-dependent dampers depend only on the instantaneous differential velocity between the damper ends. Therefore, the forces are usually out of phase with the structural shaking forces. During excitation, the maximum damper force is coincident with the maximum velocity while the structural deflection stresses are at their minimum value. On the contrary, the damping force goes to zero at maximum displacement when the structural deflection stresses are at their maximum value. Consequently, the energy dissipated is related with the relative velocity between the damper ends. Motion-dependent dampers, the third type of passive dissipation systems, depend on the vibration of a secondary system, which is usually tuned to be in resonance with the main structure. In such a way, the secondary system is oscillating out of phase with the main system, thus reducing the effect of the input energy. Friction dampers react with relatively constant force, although this force may change due to time and/or external input, such as environmental conditions (Klembczyk, 2009). In order to illustrate the relative output force functions of different types of dampers in sinusoidal motion (i.e., sine wave response), refer to Fig. 4.1. In the following sections, each of these supplementary damping devices is reviewed in depth. Reviews will be organized depending on the position, at which the devices are installed in a building: distributed throughout, or at discrete locations of the structure.

4.1.1 DISTRIBUTED DAMPING APPROACHES The most effective location for distributed dampers is where building motions are the largest. However, larger building motion leads to problem with occupant comfort as discussed in Chapter 3. Therefore, the design of a distributed damping system is a combined effort to amplify building motions without creating discomfort to building occupants. Distributed damping can be beneficial for both primary and higher mode effects within a building, but it is also often the most expensive because of the number of dampers required, and because they are not exclusively focused where building motions are the largest. Distributed dampers apply damping to a structure or mechanism, by placing the dampers at multiple locations in the structure. The main benefit of distributed damping is that the system captures and dissipates the energy at its source instead of applying damping at a secondary system, located elsewhere, or at a singular location (like tuned-mass dampers (TMDs)). Another benefit of distributed damping systems is that (direct) damping can dissipate energy in many frequencies of input vibration, instead of being confined to a tuned or “dialed-in” frequency (as per mass damping systems). This is beneficial with earthquake motions which shake the structure at several periods of vibration, in addition to the building’s fundamental period of vibration (where most of the energy is within the system). They can also work even if there is an elongation of the building periods due to inelastic behavior of the structural system.

109

FIGURE 4.1 Summary of construction, hysteretic behavior, physical models, advantages, and disadvantages of passive energy-dissipation devices for seismic protection applications. Adapted from Symans, M.D., Charney, F.A., Whittaker, A.S., Constantinou, M.C., Kircher, C.A., Johnson, M.W., et al., 2008. Energy Dissipation Systems for Seismic Applications: Current Practice and Recent Developments. J. Struct. Eng., 134:1(3), 3
21.

4.1 Passive Damping Systems

Many methods exist to implement distributed damping in a structure. The underlying concept is to connect the dampers where motion will occur, such as between beam and column joints or in between floor levels which deform relative to one another in a shearing-type motion. The dampers capture these deformations and resist in both tension and compression directions with an opposing force. For these reasons, this system functions well in typical moment frames, or braced frames, or in coupled shear wall and frame-shear wall structures (e.g., coupling beams, outrigger systems, shear wall dampers, etc.). Several configurations, where damper performance is enhanced by magnifying the damper’s perceived motion, have been introduced and implemented in the design and construction of buildings during the last two decades. Table 4.3 provides the different geometrical distributed damping configurations that are mostly utilized and also summarizes their main advantages and disadvantages. All these systems require the creation of various apparatus in the form of mechanisms that facilitate the amplification of the differential motions at the dampers’ ends, thus improving the performance of the damper. However, for the assessment of the damping system, the degrading effect of the stiffness or flexibility of these apparatus, as well as amplification characteristics, should be accounted for in the overall assessment of the damper (Huang, 2004). In the following, each different geometrical system is reviewed with particular importance to geometrical amplification of damping devices. Fig. 4.2 shows a summary of the horizontal geometrical amplification factor Gah , explaining each configuration in detail. 1. Diagonal Dampers in diagonal bracing schemes are depicted in Fig. 4.3. In this orientation, the horizontal movement of the structure only allows an angular component of the full deflection to go into the damper. This takes the motion directly to the next floor level, through a strong tension/compression member. Often this diagonal bracing scheme is considered the most basic method to apply distributed damping in a structure. Since the dampers incline horizontally, the interstory drift needs to be modified to take into account the real displacement of the damper’s ends. Looking at the case of diagonal brace dampers (Fig. 4.4), if the single degreeof-freedom (SDOF) system has a horizontal displacement u, the damper deformation ud is given by: ud 5 uGah 5 u cos θ

(4.1)

where θ is the angle between the brace and the horizontal direction. Accordingly, the damper force is related to the horizontal story shear, F, as: Fd 5 F=Gah 5 F=cos θ

(4.2)

111

112

CHAPTER 4 An introduction to dynamic modification devices

Table 4.3 Distributed Damping Geometrical Configurations: Advantages/ Disadvantages Geometrical Configurations

Advantage

Disadvantage

1. Diagonal

Simple distribution along height

Limited relative deformation (limited efficiency) Significant space requirement (less architectural freedom)

2. Chevron

Full horizontal movement transfer from structure to damper

Less efficiency due to constraints of attainable stiffness under small motion Greater space requirement (less architectural freedom)

3. Toggle

Good efficiency due to amplification of small motions (about 2
3 times) Less devices with the same efficiency of diagonal configuration

More complex design and manufacturing More difficult modeling of toggle system Greater space requirement (less architectural freedom)

4. Scissor

Good efficiency/more compact than toggle system (space saving)/ simpler installation between columns with less distance

Less efficiency compared to toggle system

5. Lever

Higher efficiency than scissor system due to higher motion amplification

Greater space requirement (less architectural freedom)

6. Damped link element

High efficiency due to both horizontal and vertical motion amplification Less number of devices requirement against diagonal configuration

Greater space requirement (less architectural freedom)

7. Cross-layer interconnection

Higher efficiency than basic diagonal configuration due to achievement of larger motion from multiple floors

Greater space requirement (less architectural freedom)

8. Outrigger connection

More suitable for tall buildings due to higher flexural deformations Increase in damping level and decrease in effective dynamic forces

More tension and compression on opposing outer columns of building Not suitable for structures with major shear deformation (low-rise buildings) (Continued)

4.1 Passive Damping Systems

Table 4.3 Distributed Damping Geometrical Configurations: Advantages/ Disadvantages Continued Geometrical Configurations

Advantage

Disadvantage

9. Wall dampers

Full horizontal shear movement transfer from structure to damper Simpler accommodation compared to diagonal configuration Simple replacement for retrofit purposes

Significant space requirement (less architectural freedom)

10. Coupling dampers

More suitable for tall buildings due to higher flexural deformations Useful for extreme seismic events by including fuse elements Suitable for architectural intention (saving space) Simple replacement for retrofit purposes

Not suitable for structures with dominant shear deformation (low-rise buildings) Exclusively useful in coupled shear walls systems

11. VE damper with cables

Significant reduction in size of damper against the conventional installation schemes (e.g., diagonal or chevron braces)

Slackness problem due to tension in cables

When designing a building under earthquake loads, one of the first assumptions typically made is that the structure has primarily shear deformations. However, when the systems grow in height, this assumption becomes more inaccurate, as the flexural deformations become comparable with the shear deformations. Accounting for both shear and flexural deformation is very important, especially when estimating the axial deformations for inclined dampers. This can lead to potentially nonconservative errors while estimating the added equivalent viscous damping given by the damper. In this case, the flexural deformations (i.e., the vertical deformation between the damper’s ends) become comparable, if not prevalent, to the shear deformations (i.e., the horizontal deformation between the ends of the damper) (Fig. 4.5). In the case of diagonal brace dampers, the flexural contribution, in the damper’s axial deformation, is calculated as follows: ud 5 uGah 2 vGav 5 u cosθ 2 v sinθ

(4.3)

where v and Gav are the vertical displacement and the vertical geometric amplification factor, respectively.

113

114

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.2 Geometrical amplification factor of damper displacement.

4.1 Passive Damping Systems

FIGURE 4.3 Dampers in diagonal-braced frames.

FIGURE 4.4 Portal scheme with diagonal viscous damper.

FIGURE 4.5 Portal scheme with an inclined viscous damper with vertical and horizontal deformations.

115

116

CHAPTER 4 An introduction to dynamic modification devices

2. Chevron Chevron frames are depicted in Fig. 4.6. In this configuration, the dampers are placed horizontally, connecting a near-rigid frame (chevron) to the floor. The advantage of this direct damping orientation is that the horizontal flexibility of the structure injects full movement directly into the horizontal orientation of the damper. However, a small amount of motion can be lost due to the constraints of the attainable stiffness of an economical chevron frame. Similar to the diagonal configuration, the geometrical factors for the chevron frame damper system are shown in Fig. 4.2. 3. Toggle Toggle frames, as shown in Fig. 4.7, can be used as a mechanism to amplify deflections into the damper in an otherwise stiff or tiny deflection situations, creating a more efficient damping system. Thus, toggle frames provide more efficient damping solution, but do require intricately designed and manufactured custom mechanisms or systems in order to perform properly. The toggle-brace damper, which is a patent of Taylor Devices Inc., in the United States, needs to be detailed by an authorized designer or manufacturer. Under the same condition of damping parameters, it can amplify the effect of damper by about 2
3 times, compared to the diagonal solution. Different chevron configurations have been developed in the past as shown in Fig. 4.8. Tianjin International Trade Center is the first project in China to adopt this configuration and subsequently it was used in several other projects including the Millennium Place and 111 Huntington Towers in Boston, MA, the Bnei Zion Hospital in Israel, and the Kimpo Airport in South Korea (Fig. 4.9). The key problems of this configuration are toggle modeling and necessary attention to proper out-of-plane constraints, as well as the physical locations of the parts of the building they are placed upon.

FIGURE 4.6 Dampers in chevron-braced frames.

4.1 Passive Damping Systems

FIGURE 4.7 Dampers in toggle-braced frames.

FIGURE 4.8 Different types of toggle configurations.

117

118

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.9 Toggle connection damper system. Courtesy of Beijing Qitai Shock Control and Scientific Development Co., Ltd.

4. Scissor The scissor configuration, patented by Constantinou (2002) (Fig. 4.10), can also make the damper over twice as effective, compared to the diagonal scheme (Fig. 4.2). It is more compact than toggle frames, so it is easier to install on structures with a smaller distance between columns. This configuration was used in both the 200 Clarendon (former John Hancock Tower) (see Section 8.2.1.3) in Chicago and Citicorp Center in New York (see Section 8.2.1.2 in Chapter 8).

4.1 Passive Damping Systems

FIGURE 4.10 Scissor connection damper system. Courtesy of Beijing Qitai Shock Control and Scientific Development Co., Ltd.

FIGURE 4.11 Lever damper.

5. Lever Fig. 4.11 shows the lever damper proposed by Rahimian (2009) where the concept of damper motion amplification can be further exponentially increased. 6. Damped link element The damped link element (Rahimian, 1999, 2007) is an example of enhancement achieved by tapping into the differential vertical motions of the structural nodes in addition to the horizontal motions. In this concept, one end of the dampers is attached to truss or wall elements and the other end to a column or independent truss, which is not integral to the main truss system. The dampers create a dynamically coupled system where the vertical motion of the end of the dampers, in combination with the lateral interstory motion, creates significant amplification of the damper’s perceived motion and thus its

119

120

CHAPTER 4 An introduction to dynamic modification devices

performance. Fig. 4.12 graphically shows the damper’s perceived motion under lateral and vertical motion. This system was applied to the design and construction of the 57-story, 225-m-tall Torre Mayor office tower, Mexico City, Mexico (Fig. 4.13). Because of the highly efficient nature of the proposed system, only 12 dampers were required at the north and south fac¸ade of the building in response to the seismic motion in the east
west direction (Rahimian and Martinez, 1998). 7. Cross-layer interconnection The cross-layer interconnection utilizes the deflections over multiple floor levels, collects the larger motion from these levels, and passes that motion

FIGURE 4.12 Damped link element.

FIGURE 4.13 Torrey Mayor (Mexico City) damped link elements (Rahimian, 2003). Reprinted with permission from Engineering News-Record, November 2007.

4.1 Passive Damping Systems

through the energy-dissipating damper connected to major structural nodes. This concept is similar to diagonal bracing, but over a much longer span. The cross-layer connection for dampers was successfully applied on 181 Fremont Street (Fig. 4.14) in San Francisco. There are four dampers on every mega brace, and three mega braces at each elevation. Therefore, a total of 32 frictionless hermetic dampers are utilized to control the vibration of the structure under earthquake and wind conditions. 8. Outrigger connection An outrigger damping solution applicable to tall and slender buildings takes advantage of the tension and compression on the opposing outer columns of

FIGURE 4.14 San Francisco 181 Fremont Street cross-layer interconnection (Almufti et al., 2016). Reprinted with permission from STRUCTURE magazine, June 2016.

121

122

CHAPTER 4 An introduction to dynamic modification devices

the building which amplifies the movements of the central core at the location of the outriggers. Often outrigger damping can be accomplished by creating a rigid level near the top of a building that moves with the core and connects dampers between the rigid level and the outer columns of the building. The system is highly efficient because of the amplified moments achieved, but unless there are multiple outrigger levels in the building, it starts becoming less of a distributed system and more mode shape specific. One of the first examples of outrigger damping system scheme, working in concert with peripheral column, is discussed in detail in Section 8.1.2 (Chapter 8) for Two Union Square in Seattle, WA. In addition, in Section 8.1.3 (Chapter 8), another case study of damped outrigger is described for Shangri-La Place in the Philippines. This solution can be represented as a scheme shown in Fig. 4.15A and the relative simplified model of perimeter columns in series with dampers (Fig. 4.15B). An alternative outrigger configuration is to utilize VE coupling dampers (VCDs) (described later on) as shown in Fig. 4.16 (Christopoulos and Montgomery, 2013). The benefits of VE material in this configuration are that it adds both an instantaneous displacement and velocity-dependent restoring force and thus the damper provides both stiffness and damping to the structure which can be efficiently used for all lateral loads: from frequent and design level winds through maximum considered earthquakes. 9. Wall dampers There are two wall damper configurations, whereby damping is provided in shear as two floors move relative to one another shearing the damper: (1) viscous wall dampers (Fig. 4.17) and (2) VE wall dampers (Fig. 4.18).

FIGURE 4.15 Simplified model of outrigger damping.

FIGURE 4.16 VCD outrigger (Christopoulos and Montgomery, 2013). Reprinted with Permission from Kinetica Dynamics.

FIGURE 4.17 Wall viscous dampers: Dynamic Isolation System Inc. (A and B) and GERB (C).

124

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.18 VE wall damper (Montgomery, 2011; Christopoulos and Montgomery, 2013). Reprinted with Permission from Kinetica Dynamics.

•

The viscous wall damper consists of a narrow tank filled with a highly viscous fluid connected to the bottom story and a vane which is connected to the top story and immersed in the viscous fluid. As the story is sheared, the top vane moves relative to the tank and shears the material. This scheme was applied in the 240-m-high DoCoMo building, Tokyo (Wada et al., 2000). This building has two parts: a lower office part with 27 stories and an upper antenna part with 24 stories. Therein, 76 viscous wall dampers are installed along two principal directions of steel frames in the lower part of the building. Accordingly, 5% equivalent viscous damping along both directions is obtained using these dampers, and the steel frame is designed such that it behaves as elastic regime under seismic loading (Wada et al., 2000). • VE wall dampers consist of VE material bonded to steel plates, and when the floors move relative to one another, they shear the VE material adding damping to the building. This scheme was applied in the 24-story SEAVANS twin building (consisting of south and north buildings), Tokyo (Yokota et al., 1992). The main structural system of this building is steel frame, where the dampers are installed along the south building in order to mitigate both wind- and seismic-induced vibrations. The damper material, thermoplastic rubber and bitumen compound were manufactured by Showa Shell Sekiyu K.K. and Shimizu Corporation (Yokota et al., 1992). 10. VE coupling dampers Coupling dampers were proposed to address the deformation mechanisms and kinematics of tall reinforced concrete (RC) buildings whereby the main

4.1 Passive Damping Systems

structural walls deform in a cantilever fashion deforming the coupling beams in shear as shown in Fig. 4.19 (Christopoulos and Montgomery, 2013). The coupling damper is composed of VE material layers that are bonded in between steel plates, with each consecutive steel layer extending out to the opposite side and anchored into the structural walls. The dampers replace structural members, such as coupling beams and therefore do not affect the architecture of the buildings in any way. When the building deforms laterally or torsionally, the walls deform relative to one another causing differential vertical displacement within the coupling elements deforming the VE material layers in shear, and thus provide both a displacement- and velocity-dependent force which couples the walls and provides instantaneous supplemental viscous damping to the structure. When properly distributed throughout the height of the building, they provide distributed viscous damping to all lateral and torsional modes of vibration under both wind and earthquake loading. The dampers can also include structural fuse elements that are designed to activate to add further hysteretic damping to the system under extreme seismic loading. They are produced by Nippon Steel and Sumikin Engineering and utilize 3M ISD-111H material. 11. VE dampers with cables Choi and Kim (2010) proposed a novel installation scheme for VE dampers with the use of cables (Fig. 4.20). With the help of this scheme, the relative displacement (story displacement) between the cable and main structure is

FIGURE 4.19 Coupling damper (Montgomery, 2011). Reprinted with Permission from Kinetica Dynamics.

125

126

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.20 Details of VE dampers connected to a steel beam and cables. Adapted from Choi, H., Kim, J., 2010. New installation for viscoelastic dampers using cables. Canad. J. Civil Eng., 37 (9), 1201
1211.

achieved and results in significant reduction in the size of dampers against the conventional installation schemes (e.g., diagonal or chevron braces). The cables can be installed continuously or discretely between the building’s base and top (Fig. 4.21). Such a new scheme can be utilized for other types of distributed-based dampers (e.g., friction dampers), but damper efficiency may be reduced due to cable stretch.

4.1.1.1 Velocity-dependent devices Velocity-dependent dampers have an output force that is defined as a function of relative (between damper ends) velocity, displacement, and frequency. In this category, there are two different types of devices: viscous (linear and nonlinear) and viscoelastic. In the following section, the mechanical and physical properties of each type of device are reviewed in detail.

4.1.1.1.1 Fluid viscous dampers Fluid viscous dampers operate by providing a resisting force only when moving. They usually do not add stiffness to a structure, and they do not carry any static load. A fluid damper consists of a piston moving back and forth through a viscous fluid, thereby generating high pressure. This piston has custom-designed orifices that produce an optimized relationship with a pressure (force) that varies with velocity. The greater the velocity, the greater the resisting force that is produced (Klembczyk, 2014). Because fluid dampers only produce a resisting force while moving and do not provide a restoring (spring) force, energy is completely absorbed into the damper’s fluid and converted to heat. This absorbed energy is simply the summation of the damping force multiplied by the deflection. Because some fluid dampers can be designed to generate pressure . 69,000 kPa, the force and, therefore,

4.1 Passive Damping Systems

FIGURE 4.21 Installation scheme for VE dampers: (A) continuous and (B) discontinuous tendons. Adapted from Choi, H., Kim, J., 2010. New installation for viscoelastic dampers using cables. Canad. J. Civil Eng. 37 (9), 1201
1211.

the absorbed energy can be relatively high. It is this absorbed energy that significantly reduces the necessity of the structural portions of the building to absorb that energy through beam/column yielding and damage. Although this seems like a simple concept, the benefits are not often fully realized. It is not only important to know how fluid dampers absorb energy, but also when they absorb energy (Klembczyk, 2014; Christopoulos and Filiatrault, 2006). Imagine a structure moving due to a transient input. A significant response of that structure will be along its dominant natural frequency as a sine wave. As the structure moves through its initial position, the deflection stress at this moment in time is zero. Instead, the velocity is at maximum and therefore the damper is reacting with its greatest force. Conversely, as the structure reaches its peak deflection and stress farthest away from its initial position, the velocity reduces to zero and therefore the damper is reacting with zero force at that moment in time (Klembczyk, 2014).

127

128

CHAPTER 4 An introduction to dynamic modification devices

The benefit of fluid dampers is to be out of phase with the structural deflection stress, and this is not the case with elements that increase stiffness or elements that are not velocity sensitive such as friction dampers or bucklingrestrained braces (BRBs). Instead, VE dampers’ stiffness component (to be discussed later) is not out of phase with the structural deflection; while the damping component is. Furthermore, fluid dampers can be designed to be almost maintenance free and maintain their full function even after a significant seismic event (Klembczyk, 2014). Basic design description. There are relatively few essential design elements of a fluid damper. However, the detailing of these elements varies greatly and it can, in some cases, become difficult and complex. Fig. 4.22 depicts an example of fluid damper and its parts (Taylor, 1999). The damper displayed in Fig. 4.22 is shown in its midstroke position. The cylinder is completely full of fluid including in chamber 1 and chamber 2. The piston rod is attached to the piston head. On the left end of the piston rod, there is a clevis (not labeled) that is used to attach the damper to the structure. As the damper reciprocates during a dynamic event, this clevis, the piston rod, and the piston head move as one component. All the other parts remain stationary. While subjected to a compression force, the piston moves the fluid from chamber 2 to 1, so that the resulting damper force is proportional to the pressure differential between the two chambers. As the piston head moves, the fluid on either side of it is forced through orifices in the piston head (orifices not shown in Fig. 4.22). On the left side of the cylinder, there is a seal and relative retainer to encapsulate the fluid against static and dynamic pressure. On the right side of the cylinder is a rod accumulator and the relative housing. This accumulator consists of a polymeric material with closed cell air bubbles. This accepts the displaced volume of the piston rod during stroking in the compression direction by compressing the fluid. This ensures that high pressure does not develop. During the extension stroke, the control valve

FIGURE 4.22 Fluid damper. Adapted from Taylor, D., 1999. Buildings: design for damping. In: Proceedings of the Boston Society of Civil Engineers, BSCES Fall 1999 Lecture Series, Dynamics for Structures.

4.1 Passive Damping Systems

FIGURE 4.23 Structure diagram of fluid hermetic damper. Adapted from Taylor, D., 2002. Taylor Devices Hermetic Dampers: Description, Applications, and Design.

opens and allows the fluid to flow freely back into cylinder chamber 2 to be ready for the next compression stroke. On the right side of the cylinder, another clevis is used. Although not shown (Fig. 4.22), both clevises are typically outfitted with spherical bearings to allow some level of misalignment with the surrounding structure and to protect damper from shear and bending forces (CEN, 2009). Alternatively metal bellow seals can be utilized as in frictionless hermetic dampers (Fig. 4.23; Taylor, 2002). This allows the viscous damper to provide a bigger power while the heat can be balanced at any time and it can withstand higher internal temperature without damage, improving the durability and stability. Additionally, dampers with metal bellows are virtually friction free. Viscous dampers are normally mounted in structural locations where some elastic recentering forces are provided by the structural frame itself. There are no specific limits for the displacement capacity of hydraulic dampers, which will only depend on the length of the device. Moreover, since the damping properties are not hysteretic, but have a viscous origin, the devices do not provide any resistance to recentering capacity. In order to provide additional recentering capacity, it might be possible to prestress the device or mount it in parallel with a springlike element, the so-called spring fluid viscous damper (CEN, 2009). However, in this case, there is a significant damping reduction, while the recentering capacity at the end of the ground motion can become significant, with low expected residual displacement. Functional description. The velocity-dependent force relationship of fluid dampers is typically characterized by the following equation (Fig. 4.24): Fd 5 cv ju_ d jα sgnðu_ d Þ

(4.4)

where cv is the viscous damper constant that is independent on the motion (i.e., scalar); u_ d is the viscous damper velocity; α is referred to as the damper velocity exponent; and sgnðÞ is the sign function. This exponent can typically be set to anywhere between 0.1 (or even less) and 2.0 depending on the specific application. Dampers with an exponent of 1 are considered linear; otherwise, they are called nonlinear damper.

129

130

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.24 Force
velocity relationship for viscous dampers.

Linear viscous dampers (LVDs) are often considered since they are the least difficult to model because they are characterized by a linear force
velocity relationship (Fig. 4.24). As such, the force generated by the LVDs is out of phase with the structural system response (Fig. 3.3 (Chapter 3)). This is considered an advantage compared to hysteretic dampers, which create the maximum force at maximum displacement. Nonlinear viscous dampers (NLVDs) have the force (Fig. 4.24) that depends on the velocity with exponent, α 6¼ 1. Typically, dampers with very low exponents, that is, lower than 0.1, produce this relationship through the use of pressure relief valves (PRVs). Caution must be exercised when using this type of configuration since PRVs are often sensitive to acceleration, high frequency, and contain moving parts. The major advantage of these devices is that, in the case of velocity spike, the force produced by a damper with a low value of α can be controlled to not overload the system. For seismic application, this exponent ranges between 0.1 and 1 (most commonly in the range 0.1
0.5) (Asher et al., 1996; Rodriguez et al., 1994). Typically, low value of α (lower than 0.4) is preferred, because they dissipate a lot of energy in short time. When compared to α 5 1, dampers with α lower than 0.2 can give the same maximum displacement with a reduced maximum force transmitted to the structure, or alternatively can reduce the displacement when keeping the same maximum force. Vice versa, when wind is the only dynamic input, very low value of α can dissipate too much energy and thus increase too much temperature; this increase in temperature of the damper can be difficult to manage for input lasting days (and not seconds like an earthquake). This is the reason why for wind application usually it is preferred to use α 5 1 or higher, up to about 1.8
2.

4.1 Passive Damping Systems

Additional information about LVDs and NLVDs are provided in Appendix A. Specification requirements. Fluid dampers are adjusted by the manufacturer for each project to meet specific customer-specified parameters. The required parameters include: 1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11.

Maximum rated force (and/or maximum design velocity) Maximum design displacement Minimum safety factors to yield Minimum required useable deflection from neutral position Connection rigidity to the surrounding main structures Damping constant Damping exponent Operating temperature Maximum wind power input (if applicable) Maximum damper envelope Damping mounting configuration, including method of replacement if applicable 12. Available space for installation 13. Testing requirements The maximum rated force of the damper is usually the force expected during the maximum considered event that the device is designed for. The yield safety factor is based upon either the maximum rated force or the velocity at which this maximum force occurs. Typically, the safety factor for LVDs is 1.5
2.0, meaning that the damper will not yield when subjected to a force or velocity 150%
200% of the rated maximum (FEMA, 2000d). In alternative, EN 15129 (CEN, 2009) amplifies the design forces utilizing a reliability factor, SF, defined as follows: SF 5 ð1 1 dt Þ1:5α

(4.5)

where dt is the design reaction tolerance given by the manufacturer. It is important to note that for NLVDs, it is not necessary to use high safety factors, because it is the nonlinearity itself that controls the maximum force when the velocity is higher than the design value. The damping coefficient, damping exponent, and temperature ranges can be easily expressed on a graph, defining allowable damper performance bandwidth at any specific operating temperature (Taylor, 1999). Viscous damper testing requirements shall be adequate to ensure proper operation during its intended life. In particular, it is recommended to perform tests on the full-scale device up to the maximum design velocity. When specified testing does not verify all requirements, analyses shall be required to prove the necessary design for any specific application (see Chapter 7 for further details). Additional considerations Seals: The seals used in a fluid damper must be capable of a long service life; at least 25 years without requiring periodic replacement. However, the design life of most buildings is 50
100 1 years, so it means the dampers will need to be

131

132

CHAPTER 4 An introduction to dynamic modification devices

replaced. The seal materials must be carefully chosen for the service life requirement and for compatibility with the damping fluid. Since dampers in structures are often subject to long periods of infrequent use, seals must not exhibit longterm sticking nor allow slow seepage of fluid. Most dampers use dynamic seals at the piston
rod interface, and static seals where the end caps or seal retainers are attached to the cylinder. For static seals, conventional elastomer O-ring seals have proven to be effective. Dynamic seals for piston rod should be manufactured from high-strength structural polymers, to eliminate sticking or compression set during long periods of inactivity. Moreover, structural polymers do not age, degrade, or cold flow over time. In comparison, conventional elastomers will require periodic replacement if used as dynamic seals in a damper (Taylor, 1999). Fluid dampers are essentially fluid-filled mechanisms which must be capable of extremely long-term service without maintenance. In addition to requiring materials that are inherently resistant to corrosion, damper materials have additional constraints, including low notch sensitivity, freedom from stress cracking, and a high impact resistance. This is especially true for the cylinder of the damper, which is a pressure vessel, and must accept substantial triaxial stresses (Taylor, 1999). A properly designed and manufactured damper should not require any type of periodic service. This is simply because, if the proper seal is selected by the manufacturer, the damper will be essentially dry sealed, with high seal scraping forces used to eliminate any static seal seepage. In the early years of dampers, most seals were produced for use in hydraulic systems, where hydraulic cylinders are used to perform work. Since a hydraulic cylinder is expected to promptly and accurately move to specific positions, even small amounts of seal friction will degrade the resolution of the hydraulic system. Thus, most hydraulic systems use dynamic seals that are intended to prevent leak, both statically and dynamically. Since dampers are passive elements, system resolution is not a design parameter, so each damper manufacturer has developed proprietary dry seals to prevent any measurable leakage during service. The types of dynamic seals used in dampers are limited by the life span of the seal, as wear occurs over time as the piston rod moves back and forth. In general, seal life is determined by measuring rod displacement, during a damper’s lifetime. Present-day seal designs are considered as so robust that a well-built damper should be warranted by the manufacturer for at least 25 years. The biggest problem encountered with viscous dampers used in TMDs (Section 4.1.2) is that they must adjust long distances, almost continually, for the life of the building. This naturally brings up the issue of seal wear, which in itself is compounded by the TMD requiring very low seal friction to allow the mass to move freely. The end result is that a conventional damper modified with low tension seals and used in a tuned-mass system will require that seals be replaced at specified intervals. This may also be true for dampers that are designed to mitigate wind effects for tall buildings. An attractive alternate to this problem is to use a damper that does not have sliding seals. To achieve this design configuration, all points in the damper where

4.1 Passive Damping Systems

a sliding seal would be placed are replaced with roller bearings. Next, flexible metal bellows are attached over the piston rod to retain fluid. The use of flexible metal bellows seals comes from the U.S. space program, where the use of hydraulic components in the vacuum environment of outer space requires that fluid be hermetically sealed at all times, to prevent outgassing (in outer space, fluid materials exposed to vacuum turn into a gaseous cloud, which can interfere with the operation of electronics or optics). Flexible metal bellows are made from multiple leaves of stainless steel sheets, edge welded to make a bellows-like shape. Any desired damper stroke can be accommodated, simply by welding on more leaves. Because the metal bellows seals by flexing, a damper using this type of seal has highly desirable near-zero friction, combined with an infinite life. The only disadvantage is that the metal bellows are extremely labor intensive and therefore costly (Taylor, 1999). Power dissipation: It is necessary for damper manufacturers to calculate the thermal response inside a damper to prevent overheating of internal parts during use. In most cases, overheating damage manifests itself by leakage, usually caused by a softened or melted dynamic seal. If calculations indicate that overheating is an issue, then, in most cases, the physical envelope of the damper will be increased in size until temperature rise during operation is low enough to be safely accepted by the internal parts. Thermodynamics teaches that there are three heat transfer processes, defined as convection, conduction, and radiation. In most cases for tall buildings, the heat transfer process will be a combination of all three. It is very important that the correct transfer processes are known for a given application when the damper is being sized, and this is the responsibility of the damper manufacturer (Taylor, 1999). Types of devices and manufacturers. In the market, there are a great quantity of types of viscous dampers and relative manufactures available. Specific technical data for each device are proprietary and manufactures do not provide any fabrication or detail for the specific design of each device. In this case, structural engineers, architects, or constructors should contact the selected manufacturers for the mechanical design of the damping devices itself. A selection of the major manufactures and the types of viscous dampers that they provide are the following: •

• •

AGOM (Italy): AGOM antiseismic hydraulic devices, called H-Safe, are composed of a cylinder filled by silicon fluid with a movable piston. The material utilized for the piston is either stainless steel or high-strength carbon steel, which may be chrome plated for seal compatibility. All other structure parts (cylinder, connections, etc.) are made in steel (S355JR and/or S275JR (CEN, 2005)). The reference European Standard Code is EN15129 (CEN, 2009). Brant Hydraulics Corporation (Taiwan): Viscous dampers are produced by this company for civil application. Dorman Long Technology (United Kingdom): Their viscous dampers use silicone oil as the damping medium which gives very stable and durable

133

134

CHAPTER 4 An introduction to dynamic modification devices

• •

•

•

•

•

•

•

performance and is rated for use in temperatures from 250 C to 250 C. The load resistance is between 60 and 3000 kN with a maximum stroke of 6 500 mm. Douce Hydro (France): Douce Hydro produces fluid viscous dampers with Zero Leakage Technology, with a velocity coefficient from 0.15 to 0.20. Dynamic Isolation System (United States): Dynamic Isolation System produces viscous wall dampers that can have two vanes. A double-vane system can also be utilized to provide twice the damping force with a similar section dimension. The fluid has a viscosity of 9000 Pa-s. FIP Industriale (Italy): FIP produces two different categories of devices: OP/ OTP (fluid viscous dampers) and OVE/OVE (spring fluid viscous dampers). The first category is a classical viscous damper, while the second one is a classical viscous damper with spring effect in parallel and—if requested—a preloaded force, which could be useful to avoid displacement under service load conditions. The classical viscous dampers (OTP series) are the most used in buildings, while spring fluid viscous dampers are sometimes used in bridges. The OP series is a device that behaves as viscous damper only above a certain force; this is typically used in bridges to keep service horizontal forces at the fixed pier. Usually the exponent is lower than 0.1 for earthquake range of velocity and 1 or more for the wind range of velocity. Freyssinet (France): Freyssinet produces a family of hydraulic dampers called TRANSPEC SHA. This type of damper has an internal compensation chamber that permits to take into account the variation in oil volume due to temperature. Moreover, it produces Isosism FD that is a classical viscous damper device filled with a silicon fluid. The utilized silicone fluid has a low compressibility (EN $ 2500 MPa) in order to prevent the storage of a portion of elastic energy. Furthermore, a proportional valve technology permits the selection of a custom constitutive law for the device. The seals are made of synthetic material with high wear resistance that are able to withstand high temperatures. Adequate corrosion protection is also provided in accordance with EN1337 (CEN, 2004) and the reference European Standard Code is EN15129 (CEN, 2009). GERB Schwingungsisolierungen Gmbh & Co (Germany): It produces viscous wall damper that consists of a steel structure connected to the lower floor and a piston to the upper floor and a viscous liquid between the two parts. Typical device dimensions are 1.5 3 0.3 3 1.1 m3 (length/width/height). This product is maintenance and abrasion free. ITT Endine (United States): Supplier of viscous dampers for both seismic and wind applications; their viscous damping devices (VDDs) are in many wind TMDs. Jinan Baidungs Machinery Equipment Co., Ltd (China): B. Don is a viscous damper that is thermostatically controlled and is virtually unaffected by temperatures from 250 C to 170 C. The materials utilized are elastomer and stainless steel piston rods for all the models. Maurer (Germany): The Maurer hydraulic damper can be utilized for displacements greater than 6 10 mm. At low velocity (,0.1 mm/s), a low force (FT , 2% of the maximum force, FI) is exacerbated and the fluid flows

4.1 Passive Damping Systems

•

•

from one side to the other. At higher velocities (.0.7 mm/s), the force is capped to a maximum limit (FI). This force is independent from frequency. The suggested damping velocity exponent is between 0.015 and 0.4 and the admissible maximum velocity is 1500 mm/s. The design service life of the device is 40 years. The range of operation is between 240 C and 180 C with tolerances of response force less than 6 5%. Oiles (Japan): The Oiles viscous wall damper is a vibration attenuator using the shear resistance force of a highly viscous fluid. It consists of an outer steel container filled with a viscous fluid and an internal shearing plate. The viscous fluid utilized has high viscosity properties and it is really stable under cycling shearing without degradation. Taylor Devices (United States): Taylor took Cold War technology from military/aerospace applications and converted that technology for use in structural dampers. All Taylor Fluid Viscous Dampers utilize solid stainless steel piston rods and are Teflon impregnated. For long-stroke applications, the piston rod is protected against bending by a heavy walled external guide sleeve. The cylinder, end cap, and sleeve are made of alloy steel and they are corrosion-protected by either painting or chrome plating. Seals run dry and are maintenance free. Taylor Devices uses one quality system conforming to AS9100 (AS, 2016) that is required for use in US military and aerospace applications. This same quality system (which meets and exceeds ISO 9001, 2015) is used for all Taylor Devices’ structural dampers. One of the major devices developed by Taylor is the hermetic damper with metal bellows seals. This was developed based on NASA and US military research that needed to operate in extreme temperatures and in a vacuum of space. The solution was to use a flexural metal bellows seal, thus sealing by nonsliding methods. The metal bellows seal consists of thin disks of stainless steel welded together to form a bellows configuration. To assure zero leakage, each bellows assembly is placed in a vacuum chamber filled with helium gas and mass spectrometer tested for acceptance. Tests demonstrate virtually friction-free operating. Hermetic dampers are reliable over a wide frequency range (0
500 Hz) (Taylor, 2002). Additionally, since the combined stresses of the metal bellows seals are held below the fatigue endurance limit, the metal bellows dampers have unlimited cycle life. Other products include dampers with special pressure relief valves (PRVs). These allow the damper to have different performances depending on the loads. With the relief valve closed, the damper performs as either a damper or a lockup device used mainly for wind excitation. When a higher velocity event occurs such as an earthquake, the PRVs will begin to open, thereby limiting the transmitted force and absorbing a specified amount of energy according to a different damping equation. After the earthquake, the damper returns to its initial state. Thus, a unique device that is optimized for both wind and seismic excitation can be provided. Moreover, Taylor provides spring fluid viscous dampers. Since the elements are separate, the amount of damping and spring combinations are

135

136

CHAPTER 4 An introduction to dynamic modification devices

virtually limitless. Large amounts of damping can be obtained in a singular device that provides spring (restoring) force. Springs are manufactured from special polymeric elements, coil springs, machined springs, or high-pressure liquid springs.

4.1.1.1.2 Viscoelastic dampers VE dampers comprise solid VE material sheets which are rigidly bonded to steel plates (Fig. 4.25) which are anchored to the lateral load resisting structure and therefore the VE material is instantaneously engaged in shear under lateral loading. In general, there are two types of VE material that are commercially available for civil applications: (1) VE material and (2) rubber. The VE material is produced by 3M and was the first damping system used in buildings in 1969 (Mahmoodi et al., 1987). It has been used in over 250 buildings for both wind and seismic vibration mitigation. The 3M ISD series of material has a very highdamping pure VE response (summation of an elastic and viscous restoring force—see Fig. 4.26A) with mechanical properties essentially consistent for all levels of deformations even at the micromillimeter level for a given frequency (Ooki et al., 2004). The second is a high-hardness rubber, which was first introduced in building structures in 2006 (Christopoulos and Montgomery, 2013). It has an elastoplastic restoring force (see Fig. 4.26B). It is also engaged

FIGURE 4.25 Example of VE device.

FIGURE 4.26 Types of VE damping material: 3M ISD material (A) and high-hardness rubber (B).

4.1 Passive Damping Systems

instantaneously but has a different elastoplastic restoring force (strain dependency) for different deformation levels. As both materials are instantaneously engaged in shear, these materials add damping to the structure in proportion to their hysteretic response for all loading types: frequent winds, ultimate level winds, service level earthquakes, design earthquakes, and maximum considered earthquakes. This strain independency is a unique feature for the VE material which gives it the credibility to add damping at extremely small deformation levels such as frequent wind events. The performance of VE dampers under such events should be confirmed by considering upper bound and lower bound of material temperature under frequencies and strains of interest (Christopoulos and Montgomery, 2013; Montgomery and Christopoulos, 2015). For the serviceability limit state design, determination of the static and dynamic loading proportion to the wind load must be done with the help of wind tunnel professionals, and the relevant static and dynamic responses should be evaluated separately (Christopoulos and Montgomery, 2013). Since frequent wind events usually prolong within a longer time, both of the VE material types described earlier should have very high durability to aging and fatigue and therefore do not require a maintenance or monitoring program to ensure adequate performance over the life of the structure. Functional description. Contrary to viscous dampers, which provide only a velocity-dependent force, VE dampers provide also a displacement-dependent elastic restoring force. The VE material behavior can be modeled using a simple Kelvin solid model which consists of spring and dashpot (in parallel) (Fig. 4.27) (Christopoulos and Filiatrault, 2006). The force
displacement hysteretic response of a VE damper (which is a series of steel plates and VE material layers combined together in the various configurations described earlier) can be defined, considering a VE material layer with a shear thickness, hs , and total shear area for all VE layers, As , as: Fd;ve ðtÞ 5 kve ud ðtÞ 1 cve u_ d ðtÞ

(4.6)

where the damper stiffness and damping coefficients read, respectively, kve 5

FIGURE 4.27 Kelvin solid model.

GE As hs

(4.7)

137

138

CHAPTER 4 An introduction to dynamic modification devices

cve 5

GC As hs

(4.8)

where GE is the elastic shear modulus and GC is the shear viscous damping constant. Given the VE material properties (hs , GE , and kve ) at the design frequency and design temperature, the shear (bonded) area of damper (As ) can be calculated by Eq. (4.7). Additional considerations regarding the mechanical and physical behavior of VE dampers are given in Appendix A.2. Material behavior. From the earlier discussion, it becomes clear that major factors for designing VE dampers are the properties of the shear storage moduli, shear loss moduli, and the loss factors. The moduli depend on different parameters, such as excitation frequency, shear strain level, and the variation in the material temperature. In the 1990s, a lot of studies were carried out in order to understand this dependency (see Chang et al. (1991), Tsai (1994), and Samali and Kwok (1995) for further details). The majority of these studies were conducted on VE materials which have been discontinued for structural applications. In general, the experimental studies showed that VE material shear modulus and loss factor are relatively constant over reasonable strain levels and increase under reduced temperature and increased frequency. In general, VE material suppliers will provide the VE material properties for the design conditions under investigation. 3M ISD-111H material. Current commercially available structural dampers use 3M ISD111 and ISD111H VE materials. The ISD111H material was specifically developed for high-rise building applications with a higher stiffness, better durability, and less temperature dependence than the previously tested 3M compounds, ISD109 and ISD110. Fig. 8.93 (Chapter 8) shows a full-scale VCD tested uniaxially in a dynamic MTS machine at the University of Toronto. The test was conducted at a harmonic frequency of 0.15 Hz (a common tall building frequency of vibration) and progressively increasing strain amplitudes. This test showed essentially strainindependent VE material properties from displacement amplitudes of 6 0.003 to 6 9 mm. This demonstrates that VE dampers using 3M VE material are one of the damper types that can robustly add viscous damping from frequent wind vibrations through maximum considered earthquake vibrations. Previous tests on other VE material, for low-rise applications, have shown self-heating during seismic excitation that can cause variation in the VE material properties during seismic loading. In high-rise buildings, the natural period of vibration is relatively long and the ISD111H is significantly less temperature dependent compared to the previous damper compounds. Fig. 4.28 shows three tests of another VCD in a uniaxial configuration with embedded thermocouples in the VE material layers to monitor the VE material temperature for the three tests. Three separate damper force
displacement hysteretic test results based on displacement input of a damper modeled in a VCD configuration in Perform-3D in an 85-story building subject to three different earthquake loads are shown. As can be seen, there is insignificant temperature change in the VE material over the three

4.1 Passive Damping Systems

FIGURE 4.28 Tests of full-scale VCDs in a uniaxial configuration subject to simulated earthquake loads on a 85-story building. Reprinted with Permission from Kinetica Dynamics.

different tests, meaning that the VE material properties are expected to be consistent over the duration of the tests. Montgomery and Christopoulos (2015) have shown that for tall building applications, self-heating is not expected to result in significant change in VE material properties for wind or earthquake loading. VE material manufacturers can provide further information for all types of loading conditions based on VE material tests. High-damping rubber. The hysteretic behavior of this damper is illustrated in Fig. 4.29, which is similar to a high-damping rubber bearing (HDRB). Some important parameters, like equivalent shear modulus, Geq , and equivalent damping ratio, ζ eq , can be defined. ζ eq can be calculated by Eq. (3.60), where the energy dissipated per cycle (ED ) is the area under the relevant hysteretic loop and the available strain energy (ES0 ) is keff u2d;max =2. Damper modeling description. Several numerical models are available for modeling VE dampers. Simpler models are widely available in commercial finiteelement packages and are usually sufficient for design or for the evaluation of global damping of the structure. More sophisticated models are also available, sometimes by the manufacturer, for detailed analyses of VE dampers. These

139

140

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.29 High-damping rubber hysteresis loop definition.

models are typically more computationally demanding and more complex. Designers should consult experts when using these models. For a background of the different models available, see Kasai et al. (1993), Fan (1998), or Montgomery (2011). Typically for the purpose of design, VE material manufacturers provide upper and lower bound VE material properties based on the expected conditions (frequency, temperature, and strain) to design the structure, and a simple Kelvin or elastoplastic model can be used. It is advisable to contact VE material suppliers and obtain VE material properties based on full-scale and VE material test results. Specification requirements. Similar to the case of viscous dampers, the structural engineers should specify the following parameters to the manufacturers for VE dampers as discussed in the upcoming chapter: 1. 2. 3. 4. 5. 6.

Maximum rated force Maximum design strain and/or displacement Operating temperature Damping mounting configuration Space available for installation Testing requirements

Type of devices and manufacturers. In the market, there are many VE dampers and relative manufacturers available. Specific technical data for each device are proprietary and manufacturers do not provide any fabrication methods or further details for the specific design of each device. In this case, structural engineers, architects, or constructors should contact the selected manufactures for the mechanical design of the damping devices itself. A selection of the major manufactures and the types of VE dampers that they provide are the following: •

3M (worldwide): 3M is the largest VE damper producer in the world, for building applications, with over 40,000 dampers sold on over 250 projects.

4.1 Passive Damping Systems

• •

•

•

•

Currently, there are two ISD series VE material types available for building projects, ISD-111 and ISD-111H. In the past, they produced ISD-110 and ISD-109 material, which was the material used in the first structural damping implementation in 1969. FIP Industriale (Italy): FIP supplies VE dampers based on high-damping rubber compounds. Kinetica (Canada): Kinetica developed the VCDs (coupling dampers) for tall RC buildings. The system is adaptable to different structural systems, such as coupling beams, core walls, and outriggers. The dampers are manufactured by Nippon Steel and Sumikin Engineering Co. and uses 3M VE damping material. Nippon Steel and Sumikin Engineering (Japan): Nippon Steel and Sumikin Engineering produce VE damper products using the 3M VE material. They currently produce VCDs (coupling dampers), axial dampers, and wall dampers. Sumitomo Riko Company Limited (Japan): Sumitomo Riko Company Limited developed wall- and brace-type dampers equipped with high-performance, high-damping rubber. The rubber has excellent damping performance and resistance to temperature variations. VSL International Ltd. (worldwide): VSL International Ltd. supplies a variety of VE damping devices using high-damping rubber. Devices consist of walland brace-type dampers for applications in buildings and can be adapted to the specific needs of different building types and vibration characteristics for new build and retrofit. Solutions include prefabricated steel shear panels, unidirectional steel struts, and K-type braces with damping nodes as well as special solutions for other building elements.

4.1.1.2 Displacement-dependent devices The history of this type of dampers dates back in the 1960s, where the plastic deformation of steel beams was usually utilized to obtain required damping for resisting building structures subjected to seismic loading (Popov, 1966). In 1968, displacement-dependent dampers made of steel (steel-beam dampers) were developed principally by Kelly in the Engineering seismology section of the physics and engineering laboratory (DSIR, Lower Hutt, New Zealand) (Skinner et al., 1993). These dampers showed much better behavior (e.g., better fatigue resistance) compared to steel structural beams. There are several references describing steel-beam dampers (Kelly et al., 1972; Skinner et al., 1974, 1975a,b; Tyler, and Skinner, 1977; Tyler, 1978; Cousins et al., 1991). Such dampers are classified into three main categories (Skinner et al., 1993): twisting-beam dampers, taperedbeam dampers, and uniform-moment dampers. The first two categories were mainly used in base-isolation systems and the third one was employed in superstructure of based-isolated bridges. To improve the seismic performance of bridge structures and to add an extra damping, incorporating some hysteretic energy

141

142

CHAPTER 4 An introduction to dynamic modification devices

dissipating devices which operate on cyclic flexural or torsional yielding of steel or on extrusion or shear of lead (Blakeley, 1979). In this section, displacement-dependent devices applicable in tall building structures are reviewed. Among all the different solutions available, the following classification can be adopted: •

•

•

•

Materials (see Section 4.1.1.2.2) • Steel • Lead • Copper • Aluminum alloys • Shape memory alloys (SMAs) Standard devices (see Section 4.1.1.2.3) • Steel dampers (added damping–added stiffness, ADAS, triangular added damping–added stiffness, TADAS, rhombic, honeycomb, dual function metallic, steel slit) • Circular plate damper • U-shaped metallic damper • Yielding steel bracing system • Knee bracing systems • Scorpion-yielding connector • Lead-extrusion devices Special connecting devices (see Section 4.1.1.2.4) • Friction dampers • Self-centering systems • Using construction weight Special structural members (see Section 4.1.1.2.5) • Shear panel damper (SPD) (or infilled steel shear wall) • Buckling restrained dampers • Tube-in-tube damper

4.1.1.2.1 Introduction to hysteretic models of displacementdependent devices The hysteretic behavior (force
displacement) of displacement-dependent devices may be different depending upon their type and material. These relationships are usually calibrated experimentally based on the different device characteristics. According to De la Llera et al. (2004), energy-dissipation devices should have a stable dissipation capacity and a representative model of its cyclic behavior. For these reasons, several models have been proposed in the past for different dissipation devices, such as: •

Elastic-perfectly-plastic model, Fig. 4.30, which does not show any strain hardening after yielding. This is usually adapted for friction-sliding devices with no initial elastic stiffness (as shown in detail in later sections).

4.1 Passive Damping Systems

FIGURE 4.30 Displacement-dependent dampers: elastic-perfectly-plastic model.

FIGURE 4.31 Displacement-dependent dampers: Bilinear model.

•

•

• •

Bilinear model, Fig. 4.31, which takes into consideration the strain hardening after the yield point. This type of behavior can be utilized to represent devices that have metallic yielding (e.g., ADAS). Bouc
Wen model, Fig. 4.32, which permits to define a more complete hysteresis loop that takes into account several factors (e.g., pinching, stiffness and strength degradation, etc.). Ramberg
Osgood model, Fig. 4.33, describes the stress
strain relations of softening type. The self-centering system, Fig. 4.34, allows the structure to go back to zerodisplacement and zero-force point at the end of every loading cycle.

See Appendix A.3 for a detailed description of the behavior of structures with hysteretic dampers.

143

144

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.32 Displacement-dependent dampers: Bouc
Wen model. Adapted from Carr, A., 2007. Ruaumoko Theory Manual. University of Canterbury and Wen, Y.K. 1976. Method for random vibration of hysteretic systems. J. Eng. Mech. 102 (2), 249
263.

FIGURE 4.33 Displacement-dependent dampers: Ramberg
Osgood model. Adapted from Carr, A., 2007. Ruaumoko Theory Manual. University of Canterbury.

4.1 Passive Damping Systems

FIGURE 4.34 Displacement-dependent dampers: Idealized displacement of a self-centering system. Adapted from Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia.

4.1.1.2.2 Materials Yielding damper properties directly correlate to properties of the main material they are composed from. These properties can include stress
strain behavior, yielding point, ductility, and toughness. The stress
strain curve of various material varies and can be obtained based on the measurement of strain (deformation) at different values of tensile or compressive stress. According to this curve, important features (e.g., modulus of elasticity, yielding point, and ductility) of the relevant material are revealed. Having determined the stress
strain curve, the yielding point (yield stress or strength) can be obtained where the material starts deforming plastically, that is, where nonlinear deformation begins. The material behaves elastically prior to the yield point. Moreover, the stress
strain curve exhibits another important material property called ductility which indicates the ability of material to be deformed plastically, that is, the percent elongation of material from yielding point to failure point (ultimate stress sustained by material prior to failure). Also, toughness, the energy per unit volume of the material that can be absorbed before rupturing, is a material property defined as the total area under the stress
strain curve.

145

146

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.35 Cyclic stress
strain hysteresis of steel elements. Adapted from Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia.

Generally speaking, yielding damper category is made of two main materials: steel and lead. These materials are reviewed in this section; moreover, other material utilized like copper, aluminum, and SMAs will be briefly reviewed. Steel. Steel exhibits an elastoplastic behavior with strain hardening after a yielding plateau. This behavior is shown in Fig. 4.35 for monotonically increasing stress (bold line) and for a cyclic reversal load. When the cycle reverses, the unloading stiffness is the same as the initial elastic stiffness. However, when the material yields and the cycle reverses, a softening is happening that is called the Bauschinger effect (Bannantine et al., 1990). This effect is greatly pronounced when the material exceeds the yield plateau (Fig. 4.35). In the cycle’s range, the material does not have a yield plateau anymore and there is a postyield stiffness. This complex hysteretic behavior is usually simplified by theoretical elastic-perfectly-plastic (Fig. 4.30) or bilinear elastoplastic models (Fig. 4.31). Usually, the hysteretic behavior of steel is divided into three parts: the skeleton, the Bauschinger curve, and the elastic unloading (Fig. 4.36). The dissipation energy is composed of the sum of the skeleton and Bauschinger curve (usually . 80% of the total plastic energy; Teruna et al., 2015). Steel behavior dissipates a lot of energy at large inelastic cycles, but it has been proven that there is a strain rate influence on the tensile behavior of mild steel (Monjoine, 1994). Monjoine (1994) has shown that the yield stress increases with an increasing strain rate, which is more pronounced in high-strength steel (Restrepo-Posada, 1993; Restrepo-Posada et al., 1994). However, this does not

4.1 Passive Damping Systems

FIGURE 4.36 Hysteresis loop decomposition. Adapted from Teruna, D.R., Majid, T.A., and Budiono, B., 2015. Experimental study of hysteretic steel damper for energy dissipation capacity. Adv. Civil Eng. 2015.

147

148

CHAPTER 4 An introduction to dynamic modification devices

have an effect on elastic and the postyield stiffness (Wakabayashi et al., 1984). For these reasons, steel elements have been widely utilized for damper devices because of their reliable and efficient source of energy dissipation when they are protected against corrosion and buckling. Lead. The other main metallic hysteretic material, lead, is mainly utilized in extrusion devices (Fig. 4.58). As a consequence, its hysteretic properties need to be derived based on the extrusion deformation of the material. The energy dissipation is created by heat and material deformation mainly due to three processes: recovery, recrystallization, and grain growth. The hysteretic behavior is similar to that of steel as shown in Fig. 4.35, with the following material and configuration characteristics (Christopoulos and FIliatrault, 2006): • • • • •

No effect on the number of load cycles No effect from environmental factors Major fatigue concerns since lead can be worked at room temperature No effect by the strain rate No effect from aging (Fig. 4.37)

Copper. The most common type of copper utilized in dampers is electrolytic tough pitch copper C11000. This copper is usually annealed, so that while strength is reduced, it becomes more ductile (Fig. 4.38, De la Llera et al., 2004). In particular, high-purity copper has been studied by De la Llera et al. (2004) that showed how the efficiency of these devices is a flexibility of the primary structure. The hysteretic behavior of copper can be defined using a two-surface hysteretic loop for metals (Dargush and Soon, 1995) as shown experimentally in Fig. 4.39 (De la Llera et al., 2004).

FIGURE 4.37 Typical lead damper hysteresis loop with insignificant aging effects. Adapted from Robinson, W.H., Cousins, W.J., 1987. Recent development in lead dampers for base isolation. In: Pacific Conference on Earthquake Engineering, vol. 2, New Zealand.

4.1 Passive Damping Systems

FIGURE 4.38 Copper stress
strain relationship (Copper Development Association). Adapted from De la Llera, J., Esquerra, C., Almazan, J.L., 2004. Earthquake behavior of structures with copper energy dissipaters. Earthquake Eng. Struct. Dyn. 33, 329–358.

Aluminum. Aluminum gives this opportunity to obtain a family of different groups of alloys (generally eight groups according to the American Association numerical classification) with different mechanical properties (Mazzolani, 2004). Aluminum compared with steel is lighter (about one-third) and better resists against corrosion due to creation of aluminum oxide film on the surface; steel needs protection in this case. Aluminum is not prone to brittle fracture, but specific attention should be paid when high ductility is required.

149

150

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.39 Typical cyclic stress
strain relationship for copper damper. Adapted from De la Llera, J., Esquerra, C., Almazan, J.L., 2004. Earthquake behavior of structures with copper energy dissipaters. Earthquake Eng. Struct. Dyn. 33, 329
358.

FIGURE 4.40 Comparison between typical stress
strain curves for aluminum alloy and steel. Adapted from Mazzolani, F.M., 2004. Competing issues for aluminum alloys in structural engineering. Prog. Struct. Eng. Mater. 6 (4), 185
196.

The mechanical properties of aluminum and steel are compared in Fig. 4.40, showing that both materials behave linear elastically with different stiffnesses. According to Eurocode (CEN, 2007), aluminum alloys do not have a yielding plateau; in fact, usually, the stress corresponding to a 0.2% residual strain is used to define the “yielding limit.” After yielding, aluminum alloys have a continuous strain-hardening behavior and the ultimate deformation capacity is lower (around 8%
12%) than that of steel.

4.1 Passive Damping Systems

The most often applications of aluminum alloys in structural engineering fields are long-span roof systems, electrical transmission towers, staircases, temporary bridges, river bridges, moving bridges, hydraulic structures, marina facilities, and offshore superstructures. In case of aluminum dampers, applications have been found in shear panels (Irandegani et al., 2014). Shape memory alloys. SMAs are a class of metallic alloys that have the capacity of recovering their original shape after being deformed even at high deformations, typical of plastic deformations in steel. This is possible due to the transformation from one phase to another (usually from austenite to martensite) and then back, through heating (martensite) or load removal (austenite, known also as superelastic). A typical SMA behavior is shown in Fig. 4.41 in which three phases can be seen: linear austenite, phase transformation, and linear martensite. The phase transformation is characterized by very low stiffness, and when unloaded, the SMA returns to its undeformed configuration, making it suitable for recentering capabilities. SMA is usually composed of two to three different metals. The most commonly utilized type for civil application is nickel-titanium, because it is corrosion free and it has superior mechanic characteristics (Dolce and Marnetto, 2000). SMAs have been applied to several structural applications: cross-bracing cables, passive structural control dampers, steel moment connections, bridges, recentering ties for monumental structures, etc. (Sharabash and Andrawes, 2009; Indirli and Castellano, 2008).

FIGURE 4.41 Typical superelastic SMA behavior. Adapted from Sharabash, A.M., Andrawes, B.O., 2009. Application of shape memory alloy dampers in the seismic control of cable-stayed bridges. Eng. Struct. 31, 607
619.

151

152

CHAPTER 4 An introduction to dynamic modification devices

4.1.1.2.3 Standard devices Having reviewed the typical materials utilized for displacement-dependent devices, in the following the most common hysteretic devices are briefly reviewed: • • • • • • • • • • • •

Added damping
added stiffness systems (ADAS) Triangular added damping
added stiffness systems (TADAS) Rhombic Honeycomb damper (HDS) Dual function metallic damper (DFMD) Steel slit damper (SSD) Circular plate damper (CPD) U-shaped metallic Yielding steel bracing system (YBS) Knee bracing system (KBS) Scorpion-yielding connector (SYC) LEDs

Added damping
added stiffness systems. ADAS system consists of a series of steel plates that deform in flexure with fixed boundaries (Christopoulos and Filiatrault, 2006). This system, developed by Betchel Power Corporation, usually connects the top of a chevron brace with the underside of a top girder and it dissipates energy through the relative displacement between the top of the brace and the beam (Fig. 4.42).

FIGURE 4.42 ADAS damper example (Christopoulos and Filiatrault, 2006).

4.1 Passive Damping Systems

There have been different studies on the optimal shape of the steel plates, which should allow to reach the plastic moment simultaneously in each plate. Some of the studies conducted in the past are the following: • •

Whittaker et al. (1991) and Tsai et al. (1993). X-shaped and triangular plate dampers. Christopoulos and Filiatrault (2006). The authors proposed the dimensional variation of the plates for (1) constant width b0 and variable depth, d ðxÞ, and (2) constant depth d0 and variable width, bðxÞ, as follows (Fig. 4.43): rffiffiffiffiffi 2x d ðxÞ 5 d0 h 2x bðxÞ 5 b0 h

(4.9) (4.10)

It can be seen from Fig. 4.43 that the depth of the plate cannot be zero because at mid height h, while the moment is Mpo 5 0; the corresponding shear force needs to be resisted: V5

2Mpo h

(4.11)

Christopoulos and Filiatrault (2006) state that the second option is better since the variable width is much smaller than the variable depth.

FIGURE 4.43 Metallic dampers optimum geometry: (A) constant width and variable depth and (B) constant depth and variable width. Adapted from Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia.

153

154

CHAPTER 4 An introduction to dynamic modification devices

Triangular added damping
added stiffness systems. This category of dampers was developed by Tsai et al. (1993) based on the ADAS system with triangular plates (Fig. 4.44). The plates are welded to the top plate but are slotted to connect at the base plate. This would not impart any gravity load on the damper and also no rotational restrain is required at the top of the brace connecting the assemblage. Rhombic damper. ADAS and TADAS systems are influenced by welded bond and axial forces and, also, it is difficult to model their force-deformation behavior, especially when strain hardening is in the nonlinear region of the stress
strain diagram (Shih and Sung, 2005). For this reason, a rhombic plate integrated with ADAS system (Fig. 4.45), for seismic resistance, was proposed by Shih et al. (2004) and Shih and Sung (2005). The proposed device has the following characteristics: •

• •

The reaction of the fixed-end moment of the TADAS system is developed in the center of the steel plate, eliminating issues with unqualified welded bonds. The roller connection at both ends eliminates the problem of transferring axial loads. The plate can uniformly yield as shown from experimental testing (Shih and Sung, 2005).

FIGURE 4.44 TADAS device hysteretic loop and dimensional parameters. Adapted from Chang, S., Tsai, K., Chen, K., 1998. Improved time integration for seudodynamic tests, Earthquake Eng. Struct. Dyn. 27, 711
730.

4.1 Passive Damping Systems

FIGURE 4.45 Rhombic damper configuration. Adapted from Shih, M.H., Sung, W.P., 2005. A model for hysteretic behaviour of rhombic low yield strength steel added damping and stiffness. Comp. Struct. 83, 895
908.

For modeling the hysteretic behavior of this device, Shih and Sung (2005) proposed the utilization of the Barber and Wen model (Barber and Wen, 1980) that alters the standard Wen model (Wen, 1976). This model has been calibrated through experimental results and readers should refer to Shih and Sung (2005) for a detailed explanation of the modeling approach. Honeycomb damper. The honeycomb damper was originally developed by Kobori et al. (1992) and it consists of steel plates with honeycomb-shaped openings that function only to accommodate loads acting within its own plane (Fig. 4.46). There are typically four different configurations for honeycomb dampers: wall, brace, beam, and postdamper systems (see Fig. 4.47). A recent experimental study was conducted by Teruna et al. (2015) on honeycomb dampers with different geometrical inputs to understand their energydissipation capabilities. The study showed that dampers have stable hysteretic loops and adequate stiffness. Dual function metallic damper (DFMD). DFMDs were developed by Li and Li (2008). The authors did a study on different metallic damper geometrical configurations and compared their properties through experimental analyses (Fig. 4.48): •

X-shaped metallic damper: It has high initial stiffness and high-loading bearing capacity. As seen from the hysteresis plot (Fig. 4.48), there is a pinching effect in the middle, mostly due to stress concentration in the middle and corner of the damper and also because the shear deformation could be higher than the bending deformation. For these reasons, it is not considered suitable for application.

155

156

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.46 Honeycomb damper example. Adapted from Kobori, T., Miura, Y., Fukuzawa, E., Yamada, T., Arita, T., Takenaka, Y., et al., 1992. Development and application of hysteresis steel dampers. In: Proceeding of the 10th World Conference on Earthquake Engineering, Rotterdam, The Netherlands.

•

•

•

•

Double round-hole metallic damper: While having a high initial stiffness, this damper configuration is undesirable, because tests revealed cracks as soon as a small deformation (of 6 mm) is reached. For this reason, it is not considered suitable for application. Strip damper: This type of damper has high initial stiffness and a good plastic deformation, except when load is reduced because of buckling. For this reason, it is not considered suitable for applications. Single round-hole damper: The hysteretic behavior shows a good energydissipated capability together with high stiffness. For this reason, it is considered suitable for application. Double X-shaped damper: This solution overcomes the buckling problem encountered with the strip damper. Moreover, the experimental results show a large stiffness and energy-dissipation capabilities. For these reasons, it is considered suitable for application.

Steel slit damper (SSD). SSD was developed by Chan and Albermani (2008). This damper is fabricated from a standard structural wide-flange section with a number of slits cut from the web, in a Vierendeel truss arrangement, rounded at the ends (Fig. 4.49). The element is a weld-free design since the connection with the main structural members is provided by a bolted connection. The device can be installed at the top of a chevron brace (Fig. 4.49), and under small

FIGURE 4.47 Honeycomb damper configuration.

158

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.48 Metallic damper geometrical configurations and experimental hysteresis loops. Adapted from Li, G., Li, H., 2008. Earthquake-resistant design of RC frame with ‘dual functions’ metallic dampers. In: Proceedings of the 14th World Conference on Earthquake Engineering, Beijing, China.

deformations, it behaves like fix-ended beams and deforms in a double curvature. At a certain displacement, both ends of the strips form a plastic hinge; therefore, the properties of this device can be expressed as a function of the strip length, l0 , the strip width, b, and the web thickness, hw (Fig. 4.49). Assuming an

4.1 Passive Damping Systems

FIGURE 4.48 (Continued)

FIGURE 4.49 SSD configuration. Adapted from Chan, R.W.K., Albermani, F., 2008. Experimental study on steel slit damper for passive energy dissipation. Eng. Struct. 30, 1058
1066.

159

160

CHAPTER 4 An introduction to dynamic modification devices

elastic-perfectly-plastic behavior, the yield force, Fdy , can be estimated as follows (Chan and Albermani, 2008): Fdy 5

nfy hw b2 2l0

(4.12)

where n is the number of strips in the device and fy is the material yield stress. The damper stiffness, kd , assuming the strips are partially fixed at their ends can be estimated as follows: kd 5 C k ns

Ehw b3 l30

(4.13)

where E is the modulus of elasticity and C k is a stiffness coefficient calibrated from experiments. One of the first applications of these devices was done in Japan in 1996 (Wada et al., 2000). It should be remarked that the small cuts created in slit dampers could increase the risk of brittle fractures and, consequently, diminish reliable plastic behavior. Therefore, particular attention should be paid to this critical point. Lee et al. (2015) carried out an experimental study on hourglass-shaped strip dampers (HSDs) under cyclic loading. They suggested that the width of the strips should be larger at the mid-height, so that undesirable shear failure does not precede the formation of plastic hinges. The authors concluded that the shape of HSD was well designed to obtain the maximum dissipation of plastic deformation energy formed during low cycle fatigue; moreover, brittle fracture follows gradual ductile fracture accompanied by plastic deformation. Circular plate damper (CPD). The energy-dissipation capacity of the circular plate damper (CPD) is coming from the geometrical elasticity and fatigue resistance around connecting parts. The in-plane shear deformation governs the behavior and this system can be easily installed in a V-brace system (Fig. 4.50). One of the major advantages of this device is that it can resist loads along all directions acting on its plane. Its behavior was studied by Adebe et al. (2013) with several experimental campaigns. The main parameter is dc =hc , where dc is the diameter and hc is the thickness. This parameter allows categorization of the hysteretic performance, of both stiffened and unstiffened CPD (Adebe and Choi, 2014). The authors found from numerical analyses, with a constant diameter and varying thickness, when dc =hc , 23:5, the section is compact with yielding prior to buckling and strength degradation; while when 23:5 # dc =hc # 29:9, the section is compact with yielding prior to buckling but after strength degradation; and when dc =hc . 29:9 yielding after buckling and strength degradation. Moreover, adding stiffeners has no significant effect on the hysteretic behavior of the damper (Adebe and Choi, 2014). U-shaped metallic damper. This damper utilizes U-shaped cold-bended mild steel elements. The elements can be disposed in uniaxial (Fig. 4.51) and biaxial (Fig. 4.52) arrangement, depending on the primary purpose of the device. Deformations of the elements happen along the main leg direction and plastic deformations are transferred from the bent part to the straight legs, through a rolling-bending motion (Bagheri et al., 2015).

4.1 Passive Damping Systems

FIGURE 4.50 Typical CPD installation. Adapted from Adebe, D.Y., Choi, J., 2014. Analytical evaluation on hysteresis performance of circular shear panel damper. Int. J. Civil Environ. Struct. Construct. Arch. Eng. 8 (6), 744
750.

FIGURE 4.51 Uniaxial U-shaped metallic-yielding damper. Adapted from Dolce, M., Filardi, B., Marnetto, R., Nigro, D., 1996. Experimental tests and applications of a new biaxial elasto-plastic device for the passive control of structures. In: Fourth World Congress on Joint Sealants and Bearing Systems for Concrete Structures, ACI SP-164, Sacramento, CA.

The main goal in the design of U-dampers is to define the major dimensions: radius ru and cross-sectional area Au (Fig. 4.53), given a set value for the damper width and thickness. Experimental studies (Aguirre and Sanchez, 1992; Dolce et al., 1996) have been conducted to determine the damper behavior. A typical hysteretic curve is shown in Aguirre and Sanchez (1992) that shows the enclosed

161

162

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.52 Biaxial U-shaped metallic-yielding damper with circular arrangement. Adapted from Dolce, M., Filardi, B., Marnetto, R., Nigro, D., 1996. Experimental tests and applications of a new biaxial elasto-plastic device for the passive control of structures. In: Fourth World Congress on Joint Sealants and Bearing Systems for Concrete Structures, ACI SP-164, Sacramento, CA.

FIGURE 4.53 U-damper geometry. Adapted from Aguirre, M., Sa`nchez, A.R., 1992. Structural seismic damper. J. Struct. Eng. 118 (5), 1158
1171.

4.1 Passive Damping Systems

range of loop for a testing of sinusoidal cycles (20% variation between solid and dashed line). Based on these tests, optimal values for ru and Au are 4.5 and 10 cm, respectively. Moreover, the U-damper has no effective viscous behavior (no frequency dependence). Aguirre and Sanchez (1992) recommended that for material and production uniformity, strip stock should be used, and among all the possible arrangements, they proposed two possible frame configuration solutions, as shown in Fig. 4.54. The authors also recommended small U-elements because they are easier to construct and typically avoid unexpected failure.

FIGURE 4.54 Typical U-damper frame configuration. Adapted from Aguirre, M., Sa`nchez, A.R., 1992. Structural seismic damper. J. Struct. Eng. 118 (5), 1158
1171.

163

164

CHAPTER 4 An introduction to dynamic modification devices

Yielding steel bracing system (YBS). Yielding steel bracing systems (YBSs) may also be seen as an improvement on X-bracings ductility. Round bars absorbers can be placed at the intersection of the diagonals of an X-braced system (Tyler, 1985) (Fig. 4.55). The edges of the rectangle are subjected to bending moments and energy is dissipated by small plastic hinges that are formed in the corners. Knee bracing system (KBS). The ductility of concentrically braced frames can also be improved by using the knee bracing system (Fig. 4.56; Aristizabal-Ochoa, 1986; Balendra et al., 1990, 1992; D’Aniello et al., 2014). In this case, the energy-dissipation mechanism is coming from the flexural knee yielding at its ends and mid-section (Aristizabal-Ochoa, 1986). One drawback of this system is when the brace buckle in which case the hysteretic behavior will have an undesirable pinching behavior (similar to concentrically brace frame). For this reason, diagonal elements should carry only tension. Furthermore, double knee brace systems would be recommend for high lateral load reversal (AristizabalOchoa, 1986). The knee element can be placed in different locations: (1) top or bottom of a single knee brace, (2) both ends of a single knee brace, and (3) double knee brace system (Fig. 4.56). Balendra et al. (1992) have found that no significant ductility improvements were found in the double brace system. The optimal knee bracing system shape is selected based on the configurations of Fig. 4.56, from elastic analyses (Naeemi and Bozorg, 2009). The optimal angle is determined at a tangential ratio of almost one of knee  elements  bk =hk L1 =H , which means the knee elements should be parallel to the diagonal direction of the frame and the diagonal elements should pass through the knee point and the beam
column intersection as shown in Fig. 4.56 (Naeemi and Bozorg, 2009). Aristizabal-Ochoa (1986) proposed some general guidelines for knee elements based on the yield mechanism of a fixed
fixed beam under a concentrated load at the mid-span. He recommends using knee elements that are at least 50% lighter

FIGURE 4.55 Tyler’s yielding steel bracing configuration. Adapted from Tyler, R.G., 1985. Further notes on a steel energy-absorbing element for braced framework. Bull. NZ Natl. Soc. Earthquake Eng. 18 (13), 270
279.

4.1 Passive Damping Systems

FIGURE 4.56 Knee brace configurations: (A) K-configuration, (B) X-configuration, (C) single brace with one knee element and (D) with two knee elements. Adapted from Naeemi, M., Bozorg, M., 2009. Seismic performance on knee braced frame. World Acad. Sci. Eng. Technol., 26, 976
980.

than the bigger element between the column and beam size with a plastic moment capacity of less than: Fdy lk sin θk =8

(4.14)

where Fdy is the yield force capacity of the diagonal element; lk is the knee element length; and θk is the angle between diagonal and knee element. This is to guarantee that the knee yields first and the diagonal element remains elastic. A similar system was proposed by Balendra et al. (1994) in which the knee yields in shear, instead of bending. Indeed, the length of the knee will influence the stiffness of the frame and the mode of yielding; a longer knee will yield in flexure and a shorter one in shear. The limit length between these two behaviors can be computed as follows: lk , 2

Mp Vp

(4.15)

165

166

CHAPTER 4 An introduction to dynamic modification devices

  Mp 5 hkf bk dk 2 hkf fy

(4.16)

  fy Vp 5 hkw dk 2 hkf pffiffiffi 3

(4.17)

where fy , dk , hkf , bk , and hkw are the yield stress, depth, flange thickness, flange width, and web thickness of the knee member, respectively. Eq. (4.15) is satisfied for a high section modulus to the shear area ratio. Lateral buckling on the knee should also be prevented since it could affect the ductility (Balendra et al., 1994). Scorpion-yielding connector (SYC). SYCs are replaceable and hysteretic fuses installed in concentrically braced frames to improve the performance of existing or new buildings (Fig. 4.57). Installing such connectors into the diagonal bracing system is also called the cast steel-yielding brace system (Gray et al., 2012, 2014). The two specially designed cast steel connectors of this system are welded to the ends (flange) of each brace as shown in Fig. 4.57. SYC comprises an elastic arm and some yielding fingers, where the elastic arm is connected by weld to the brace system. Meanwhile, the fingers are bolted to a splice plate. Details of this kind of connector and relevant connections are illustrated in Fig. 4.57. This type of damper behaves elastically under service loading and dissipates input energy of stronger loads by flexural yielding of the fingers. The hysteretic

FIGURE 4.57 Scorpion-yielding connector. YBS, yield brace system. Adapted from Gray, M.G., Christopoulos, C., Packer, J.A., 2014. Cast steel yielding brace system for concentrically braced frames: concept development and experimental validations. J. Struct. Eng. 140 (4), 04013095.

4.1 Passive Damping Systems

behavior of such damping system is almost symmetric in tension and compression (Gray et al., 2012). The yield force, the required axial force for flexural yielding of the fingers, is given by (Gray et al., 2012): Fdy 5

nb0 hsyc fy 4Lsyc

(4.18)

where n is the number of yielding fingers; b0 is the width of yielding finger at its base; hsyc is the thickness of yielding finger; and Lsyc is the length of yielding finger. Fig. 4.57 more clearly illustrates the defined parameters. Lead-extrusion devices (LED). LEDs were invented in the 1970s by Robertson and Greenbank (1975). They consist of a thick-walled tube with a tie rod connected to a piston on each side. The lead is positioned in between the pistons and it is lubricated in order to prevent friction with the cylinder wall (Robertson and Greenbank, 1975). Each piston is then connected to the structure and the relative movement between each side of the lead is extruded back and forth through an orifice. This produces deformation of the lead, consuming the excitation energy (Sadek et al., 1996). Lead-extrusion dampers are useful devices for absorbing energy transferred to a structure from earthquake (Skinner et al., 1993). There are two major types of extrusion dampers: bulged shaft and constricted tube (Fig. 4.58). The bulged shaft is usually preferred for economical and performance reasons (Cousin and Porritt, 1993).

FIGURE 4.58 Type of LED dampers: constricted tube and bulged shaft. Adapted from Cousins, W.J., Porritt, T.E., 1993. Improvements to lead-extrusion damper technology, Bull. NZ Natl. Soc. Earthquake Eng. 26, 342
348.

167

168

CHAPTER 4 An introduction to dynamic modification devices

The high amount of resistive force that LED requires produces a great amount of energy absorption, much greater than that of an equivalently sized viscous damper (Rodgers et al., 2006). There are two limitations in the amount of energy dissipated: (1) shaft yield load that has economical and physical limitations, and (2) heating of the lead due to repeated cycles. One of the major problems in utilizing these type of dampers is the creation of a void during extrusion, due to the compression of the lead, expansion of the cylinder wall, and imperfections. This problem can be overcome if the lead is prestressed in order to minimize the formation of void and reduce the damper size (Rodgers et al., 2006). Lead dampers can be utilized in different locations: • • •

For base isolation, which will be discussed later in Section 4.2 For typical brace systems In steel and RC connections (Fig. 4.59; Rodgers et al., 2006), which allows reduction of the joint rotation (i.e., reducing the damage under seismic events)

FIGURE 4.59 Possible application of LED. Adapted from Rodgers, G.W., Chase, J.G., Mander, J.B., Leach, N.C., Denmead, C.S., 2006. Experimental development, tradeoff analysis and design implementation of high force-to-volume extrusion damper technology. Bull. NZ Natl. Soc. Earthquake Eng. 40 (2), 35
48.

4.1 Passive Damping Systems

The modeling of this type of damper is not straightforward, but Pearsons et al. (2006) utilized a classical extrusion process modeling approach that relates to the extrusion force, the damper geometrical and material properties, as follows:  Fd;led 5

  

Aled 1 fy expðχÞ 2 fy ðAled 2 aled Þ fy ln aled

(4.19)

where Fd;led is the extrusion force, fy is the yield strength of the working material, Aled is the annular area around the shaft, aled is the annular area of orifice (Fig. 4.60), and χ is a constant equal to: χ5

4μled Lled dled

(4.20)

where μled is the coefficient of friction between the lead and steel shaft, Lled is the length of the sliding part of the shaft, and dled is the effective diameter corresponding to Aled .

4.1.1.2.4 Special connection devices In this category, those devices having special connection details are briefly reviewed, such as: • • •

Friction dampers Self-centering systems Using construction weight

Friction dampers. Friction dampers dissipate energy through friction between plates that are connected by bolts placed in slotted holes, which allow

FIGURE 4.60 LED geometrical configuration. Adapted from Rodgers, G.W., Chase, J.G., Mander, J.B., Leach, N.C., Denmead, C.S., 2006. Experimental development, tradeoff analysis and design implementation of high force-to-volume extrusion damper technology. Bull. NZ Natl. Soc. Earthquake Eng. 40 (2), 35
48.

169

170

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.61 Pall friction damper. Adapted from Pall, A., Pall, R.T., 2004. Performance-based design using pall friction dampers—an economical design solution. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1
6, Paper No. 1955.

displacements. Many structural configurations are available (similar to metallic dampers), since they can be placed either on a (diagonal) member or at the intersection of braces in an X-braced system (an example is shown in Fig. 4.61 (Pall and Pall, 2004)). The major advantage of this type of device is that their behavior is quite easy to control, only by means of the friction coefficient. Moreover, they are easy to construct and inexpensive. However, maintenance and aging effects might pose some problems. Friction damper behavior is based on the standard friction theory. The friction force, Fds, is independent on the contact area and the sliding velocity (for relatively low amplitudes), but proportional to the normal force applied. Consequently, it can be calculated as Coulomb friction (Fig. 4.62): Fds 5 μs FN

(4.21)

where FN is the applied normal force and μs is the coefficient of friction. See Appendix A.4 for a detailed discussion about the damper behavior.

4.1 Passive Damping Systems

FIGURE 4.62 Coulomb friction SDOF model.

There are different types of friction surfaces that consequently have various coefficient of friction. These can be subdivided into two major categories: • •

Metal to metal (e.g., brass, cobalt; Tremblay and Stiemer, 1993) Metal to teflon (e.g., Bondonet and Filiatrault, 1997)

Based on all these considerations, the following four categories of friction dampers are discussed: • • • •

Slotted-bolted connections (Tremblay and Stiemer, 1993) Sumitomo friction devices (Aiken and Kelly, 1993) Pall friction devices (Pall and Marsh, 1982) Rotational friction dampers (Mualla and Belev, 2002)

Slotted-bolted connections Typical steel frame joints can be assembled to have slotted-bolted connections (Fig. 4.63), which slip under a certain loads, thus dissipating energy before the inelastic deformation of the principal dampers. The slot length should be enough to accommodate the maximum deformation and spring washers are usually required to maintain a constant slip load (Tremblay and Stiemer, 1993). Sumitomo friction devices The Sumitomo friction device was introduced by Sumitomo Metal Industries Ltd. in Japan (Aiken and Kelly, 1990). It is composed of a steel tube internally equipped with a precompressed spring, which, through an inner and outer wedge, applies a normal force on copper friction pads. The friction surface is between the steel tube and the pads (Fig. 4.64). Pall system The pall system was developed by Pall Dynamics Ltd. (Pall and Marsh, 1982) and it consists of slotted plates that intersect between diagonal braces in moment frame structures (Fig. 4.65). The clamp friction force is given by high-strength steel bolts and they are usually designed to slip only under severe seismic loading (not under service or moderate earthquakes).

171

172

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.63 Example of steel frame slotted-bolted joint. From Tremblay, R., Stiemer, S.F., 1993. Energy dissipation through friction bolted connections in concentrically braced steel frames. In: ATC 17-1 Seminar on Seismic Isolation, Passive Energy Dissipation, and Active Control, vol. 2, pp. 557
568.

FIGURE 4.64 Example of Sumitomo friction device. Adapted from Aiken, I.D., Kelly, J.M., 1990. Earthquake simulator testing and analytical studies of two energy-absorbing systems for multistory structures. Report No. UCB/EERC-90/03, EERC, University of California, Berkeley, CA.

One of the main features of these devices is to dissipate energy into two slip joints independently on the buckling capacity of the bracing. The relative hysteretic behavior can be separated into five different steps, as shown in Fig. 4.66 (Christopoulos and Filiatrault, 2006). Rotational friction system The rotation friction damper device (RFD) consists of several plates rotating against each other. Special circular friction pad disks are placed between the steel plates in order to have dry friction lubrication in the unit and to ensure a stable friction force (Mualla and Belev, 2001, 2017) (Fig. 4.67). Bolts connect the plates of the damper to each other. This adjustable bolt is used to control the normal force applied on the friction pad disks and the steel plates. In order to keep a constant clamping force, several disks spring washers are used.

4.1 Passive Damping Systems

FIGURE 4.65 Example of Pall system for moment-resisting frames. Adapted from Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia.

Hardened washers are placed between these springs and steel plates to prevent any marks on the steel plates from the disk springs. When a lateral external force excites a frame structure, the girder starts to displace horizontally due to this force. When the forces reach the designed capacity force, the steel plates of the dampers start to rotate and kinetic energy is converted into heat in the friction layer between friction pads and the steel plates, thus reducing the vibration of the structure (Figs. 4.68 and 4.69).

173

174

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.66 Pall friction damper hysteretic steps. Adapted from Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia.

FIGURE 4.67 RFDs, four-joint type. Courtesy of Damptech A/S.

The RFD can be arranged in different configurations of bracing systems and with multiunits based on the requirements of the designed friction force and the space limitations (Fig. 4.70). In addition, the RFD can be designed to have a multilevel capacity. For example, a two-level damper can be designed to have different capacities for lower vibration amplitudes, for example, induced by wind action or moderate earthquakes and for second level of severe vibrations caused by strong earthquakes (Fig. 4.71).

4.1 Passive Damping Systems

FIGURE 4.68 (A) Frame structure moved to the left side, (B) frame structure at initial position, and (C) frame structure moved to the right side. Courtesy of Damptech A/S.

FIGURE 4.69 (A) Damper at initial position, (B) in tension, and (C) in compression. Courtesy of Damptech A/S.

FIGURE 4.70 RFD with multiunits. Courtesy of Damptech A/S.

175

176

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.71 Two-level RFD with four joints: (A) configuration and (B) typical hysteresis loop. Courtesy of Damptech A/S.

A typical configuration of inverted V-bracing with two dampers is shown in Fig. 4.72. These dampers were installed in 2014 in the Abeno Harukas 300 project, the tallest building in Japan. Self-centering systems. During an exceptional event (especially a seismic one), residual deformation can be a relevant problem that can preclude the building utilization and increase the repair costs. Hence, a new device category, selfcentering system, has been introduced to reduce residual deformations. Typical behavior for this system can be idealized as a flag-shaped response, as shown in Fig. 4.73B (Christopoulos, 2002). The dissipated energy is reduced, compared to a standard yielding system Fig. 4.73A, but at every cycle, the system returns to a zero force
zero displacement position. The two responses can be related with two independent response parameters ε1 and ε2 (see Fig. 4.73). The coefficient ε2 reflects the energy-dissipation capacity and the following three limits can be defined: ε2 5 0 represents a bilinear elastic system, ε2 5 1 represents the upper bound of the recentering capacity, and ε2 5 2 represents bilinear elastoplastic hysteresis. Christopoulos (2004) has deeply studied the frequency response of these systems. The curves in Fig. 4.74 show an example of behavior as a function of the coefficients ε1 and ε2 , the normalized excitation amplitude, aexc , the normalized steady-state amplitude, a exc , and the excitation frequency ratio, σ (Christopoulos and Filiatrualt, 2006). As it can be seen, there is soft system behavior with a decrease in resonant frequency for an increasing amplitude of excitation. Moreover, the amplitude of the flag-shaped behavior is higher compared to the elastoplastic system (Fig. 4.74; Christopoulos and Filiatrualt, 2006).

4.1 Passive Damping Systems

FIGURE 4.72 2000-kN damper installed in the tallest building in Japan Abeno Harukas 300 (as per January 2018). Courtesy of Damptech A/S.

FIGURE 4.73 Idealized response of elastoplastic (A) and self-centering (B) structure. Adapted from Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia.

177

178

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.74 Frequency response curve: flag-shaped and elastic plastic. Adapted from Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia.

As shown in Fig. 4.74, the frequency response shows an unstable region that is highlighted in Fig. 4.75 for various normalized excitation amplitude and ε1 5 0:1 and ε2 5 0:8 (Christopoulos and Filiatrault, 2006). This shows a jump in the response that depends on how the excitation frequency varies (Fig. 4.76; Christopoulos and Filiatrault, 2006). The jump is higher when the frequency is decreased from higher to lower frequencies. The above briefly discussed behavior is usually associated with rocking walls, posttensioned concrete and steel moment
resisting frames, SMA, and some type of special viscous dampers (Christopoulos and Filiatrault, 2006). In the following, three possible utilization of this technology are briefly reviewed. Recentering spring systems Mechanical devices can be utilized to achieve recentering capacity. In this case, devices that consist of spring elements can be mentioned and categorized as tension-compression isolator (Taylor and Lee, 1987); elastomeric spring damper (Constantinou et al., 1998); energy-dissipating restraint (Nims et al., 1993) or preloading spring-friction dampers (Constantinou et al., 1998); and dampers with ring springs (Filiatrault et al., 2000). A shock isolation system, called tension-compression isolator, was developed to maintain precisely its alignment at pre- and postshock conditions. In this system, a liquid spring-damper component is adapted as a compression-acting device, and a slotted external tube and linkage were added as a mechanism under tension shock (Taylor and Lee, 1987). Fig. 4.77 illustrates a fluid-restoring device including the liquid spring-damper element (Constantinou et al., 1998).

4.1 Passive Damping Systems

FIGURE 4.75 Frequency response curve: unstable region for ε1 5 0:1 and ε2 5 0:8: Adapted from Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia.

FIGURE 4.76 Frequency response curve: jump phenomenon. Adapted from Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia.

179

180

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.77 Pressurized fluid restoring device including liquid spring-damper component. Adapted from Tsopelas, P., and Constantinou, M. C., 1994, NCEER-Taisei Corporation Research Program on Sliding Seismic Isolation Systems for Bridges: Experimental and Analytical Study of a System Consisting of Sliding Bearings and Fluid Restoring Force/Damping Devices, Technical Report NCEER-94-0014, National Center for Earthquake Engineering Research, Buffalo, NY.

An elastomeric spring damper utilizes a silicon-based elastomer giving both spring and hysteretic behavior. Pekcan et al. (1995) proposed both single-acting device (operative under only compression) and double-acting device (operative under both tension and compression). The second type is the modified version of the first type based on adding a housing around the device. Beyond the internal preload, the piston of damper causes the elastomer to compress, enforcing to flow over the orifice between piston head and inner casing. This leads to create both stiffness and damping. In the case that no external force exists, the initial internal pressure allows the damper to recenter (Pekcan et al., 1995; Constantinou et al., 1998). The first ring-spring device was originally manufactured by Fluor Daniel, Inc. and it consists of a cylindrical wall that incorporates a spring system with a friction wedge and an internal end stop to limit the motion (Fig. 4.78, Nims et al., 1993). The spring force creates a normal force on the steel/bronze wedges that is applied to the cylinder wall and this determines the slip force in the device. In fact, the difference between the spring force and slip force tends to recenter the damper to its initial position. Since the spring force is always larger than the sliding force, it causes device centering when there is no external force (Nims et al., 1993). The hysteretic behavior of such a device can be adjusted based on the spring stiffness and preload, as well as the gap length (Nims et al., 1993). The second device utilizes a ring-spring system as the main component. When the column spring is loaded in compression, the rings go in compression (inner) and tension (outer) (Fig. 4.79). A device of this type, called the SAPHIA damper, was developed by Spectrum Engineering, Canada Fig. 4.80. Posttensioned rocking systems The basic idea behind the posttensioned rocking system is to permit rocking motion with unbound post tensioning cables that allow

4.1 Passive Damping Systems

FIGURE 4.78 Energy-dissipating restraint. Adapted from Nims, D.K., Richter, P.J., Bachman, R.E., 1993. The use of energy dissipating restraint for seismic hazard mitigation. Earthquake Spectra 9 (3), 467
489.

FIGURE 4.79 Ring-spring details. Adapted from Filiatrault, A., Tremblay, R., Karr, R., 2000. Performance evaluation of friction spring seismic damper. ASCE J. Struct. Eng. 126 (4), 491
499.

recentering capabilities (PRESSS Program University of San Diego, California; Priestley, 1991). The rocking mechanisms usually considered are beams on column, segmental columns on each other and on their foundations, and walls on foundation. In most of the cases, these systems do not dissipate a great amount of energy at each cycle and, for this reason, hybrid systems were introduced in which cables

181

182

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.80 SHAPIA damper. Adapted from Filiatrault, A., Tremblay, R., Karr, R., 2000. Performance evaluation of friction spring seismic damper. ASCE J. Struct. Eng. 126 (4), 491
499.

FIGURE 4.81 Hybrid frame system.

are combined with longitudinal, nonprestressed (mild) steel that is allowed to yield (Fig. 4.81) (Stanton and Nakaki, 2002). Similar concepts can be utilized for rocking precast concrete shear walls (Stanton and Nakaki, 2002) in which walls are split in order to allow them to rock around their bases. Additional energy dissipation can be provided by connecting the wall edges with the foundation, with mild steel and/or ductile shear connectors between wall panels (Fig. 4.82). Similar to concrete structures, a comparable approach was proposed for confined masonry walls systems (Toranzo et al., 2004) and steel structures (Ricles et al., 2001). An example of a self-centering system is proposed by Christopoulos et al. (2002a,b) and it consists of a posttensioned energy-dissipating (PTED) system, in addition to dissipative ED bars (Fig. 4.83). The advantage of this system

4.1 Passive Damping Systems

FIGURE 4.82 Rocking wall system.

FIGURE 4.83 PTED damper example.

consists in its capacity to reposition the system that leads to have a zero
force
zero displacement point at every cycle of seismic loading. Using the weight of the construction. An alternative technique to dissipate energy, to those ones presented earlier, is to use of the weight of the construction to oppose the seismic movement. The building is placed on a system of cylinders with spherical ends (Fig. 4.84; Dima and Stef˘ ¸ anescu, 2003) or on a system of orthogonally arranged elliptical rolling rods (Rawat et al., 2015). Due to the spherical surfaces of the cylinders or rods, when a horizontal translation occurs, the building is lifted and the weight acts against the movement trend.

183

184

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.84 Seismic cylinder.

FIGURE 4.85 Examples of shear panel dampers: infilled wall type, wall type. Adapted from Katayama, T., Ito, S., Kamura, H., Ueki, T., Okamoto, H., 2000. Experimental study on hysteretic damper with low yield strength under dynamic loading. In: Proceedings of the 12th World Conference in Earthquake Engineering, Auckland, New Zealand; Tanaka, K., Sasaki, Y., 2000. Hysteretic performance of shear panel dampers of ultra low-yield-strength steel for seismic response control of buildings. In: Proceedings of the 12th World Conference in Earthquake Engineering, Auckland, New Zealand.

4.1.1.2.5 Special structural members In this last category reviewed, devices are categorized into special structural member types: • • •

SPD Buckling restrained brace (BRB) Tube-in-tube damper (TTD)

Shear panel damper. The SPD is composed of large steel plate shear walls, installed in the framework of a building (Fig. 4.85). The wall configuration allows the plate to be under in-plane shear and the yielding can spread over the entire plate (Katayama et al., 2000).

4.1 Passive Damping Systems

FIGURE 4.86 SPD with rib plates. Adapted from Tanaka, K., Sasaki, Y., 2000. Hysteretic performance of shear panel dampers of ultra low-yieldstrength steel for seismic response control of buildings. In: Proceedings of the 12th World Conference in Earthquake Engineering, Auckland, New Zealand.

Tanaka and Sasaki (2000) estimated the performance of a panel damper with a generalized width
thickness ratio, normalized by yield-strain and yield-ratio. They also introduced a modified width
thickness ratio in case the panel dampers have reinforced rib plates (Fig. 4.86). This parameter was originally utilized to determine the elastic buckling strength of the shear panel. Infilled steel shear walls (Fig. 4.85), consisting of slender unstiffened steel plates connected to beams and columns, show a very good capacity to dissipate energy (Elgaaly et al., 1993; Caccese et al., 1993; Driver et al., 1997; Elgaaly, 1998). They can be used effectively for earthquake-resisting systems due to their high strength for postbuckling behavior (Elgaaly et al., 1990; Park et al., 2007). That being said, elastic buckling strength is not considered synonymous with failure because the ultimate capacity of the system is many times bigger than the buckling load, in the case adequate support is situated at the boundaries (Elgaaly et al., 1993). While a significant pinching effect is present in the hysteresis response, when properly detailed, the wall strength does not drop due to the development on the tension field action and it can sustain high ductility demands. Looking at the production side, the frame to steel shear wall connection is generally welded, even though for economic reasons, bolt connections can be preferred depending on the infill plates, since the number of bolts can significantly increase (Fig. 4.87; Choi and Park, 2009). For this reason, partially connected infill plates were also investigated by Xue and Lu (1994) but were never validated. A modification on this system was proposed by Hitaka and Matsui (2003) that introduced slits in the panel, as shown in Fig. 4.88. The behavior of this configuration is similar to a series of flexural links that go under large flexural deformations. These links are very susceptible to buckling, but the behavior is stable due to small horizontal forces that each slit attracts (Hitaka and Matsui, 2003). An

185

186

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.87 Details of fish plate connecting infill plates and frame members. Adapted from Choi, I.R., Park, H.G., 2009. Steel plate shear walls with various infill plate designs. J. Struct. Eng. ASCE, 135 (7), 785
796.

FIGURE 4.88 Steel shear wall with slits. Adapted from Hitaka, T., Matsui, C., 2003. Experimental study on steel sear wall with slits. J. Struct. Eng. 129 (5), 586
595.

evolution of this system was proposed by Jin et al. (2016) and included an inclined slot pattern. In order to enhance the elastic flexural capacity (higher elastic buckling strength), steel stiffeners can be added (Jin et al., 2014). Another particular enhancement was proposed by Hitaka et al. (2009) who introduce slit shear steel plates attached to only the beams in the building and buckling is restrained with timber panels and slits (Fig. 4.89). Buckling-restrained braces (BRB). BRB was originally developed in Japan by Wakabayshi et al. (1973). This system is sometimes also referred to as an

4.1 Passive Damping Systems

FIGURE 4.89 Steel shear wall with slits and timber panel configuration. Adapted from Hitaka, T., Ito, M., Murata, Y., Nakashima, M., 2009. Seismic behavior of steel shear plates stiffened by wood panels. In: Proceedings of STESSA 2009—Behaviour of Steel Structures in Seismic Areas, Philadelphia, PA, pp. 623
628.

unbonded brace and consists of a steel core inside of a concrete member (reinforced or not) or inside of a steel jacket (Fig. 4.90). There is no bond between the core plate and the restraining part, so the axial force in the core is not transmitted to the restraining part. This is provided when, given enough clearance, an unbounded material could be used (Iwata, 2004). Through controlled buckling of the core inside the jacket, a quasisymmetrical behavior in tension and compression is achieved, contrary to a standard brace system that has an asymmetrical hysteretic behavior, since it buckles in compression (Fig. 4.91). This is prevented in a BRB system since the continuous lateral restrain strongly diminishes the buckling length of the core. Among all the advantages, BRBs have low postyield stiffness that could lead to the concentration of damage and different tensile, compressive capacity that could lead to problems in the V-brace configurations. Moreover, in the case the strength of the system is close to the design force, this might reduce the system overstrength when ductility demand is slightly increased. However, this is considered a minor drawback (Wijanto, 2012). Furthermore, it is not possible to inspect the core when an infill is used. BRBs can be designed in different geometrical configurations. Kersting et al. (2015) use five different configurations in their study: inverted V-bracing

187

188

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.90 BRB geometry and typical bracing layout.

(chevron), V-bracing, diagonal bracing (same direction), diagonal bracing (zigzag), and multistory X-bracing (Fig. 4.92). The multistory X-bracing reduced both axial and vertical unbalance loadings in frame beams, as well as distributed more efficiently the yielding across multiple stories. Instead, in V-bracing configurations, frame beams need to be designed for the unbalanced vertical force. This unbalanced load will be against gravity in the case of inverted V-bracing. Diagonal bracing is allowed by code (ASCE, 2017a) and a zig-zag configuration is preferred since it reduces the unbalanced axial load in the frame beams. BRBs are not used in a K-brace configuration, as it is not accepted in seismic structures or in concert with X-brace configurations, AISC 341 (AISC, 2016a). BRBs can be utilized as a knee brace (Kim and Seo, 2003), especially in the case of retrofit solutions (Fig. 4.93; Shin et al., 2012). BRB behavior is a function of material and component configuration. The most critical are: • • • •

Cross section Encasing tube Filler material Debonding agents

4.1 Passive Damping Systems

FIGURE 4.91 Behavior of conventional brace and BRB. Adapted from Xie, Q., 2005. State of the art of buckling-restrained braces in Asia. J. Construct. Steel Res. 61, 727
748.

FIGURE 4.92 Comparison of different types of BRB configurations.

In literature, there are several cross-sectional configurations available for BRBs. The most common are the cruciform and flat plate shapes (Fig. 4.94; Xie, 2005), such as (A) plate brace with mortar-infilled steel tubes, (B) H-section steel brace enclosed by RC, (C) crisscross cross-section steel brace with steel-fiber RC, (D) steel plate brace stiffened by two bolt-connected precast concrete panels, (E) wide-flange section with a restrained steel tube, (F) two circular steel tubes, (G)

189

190

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.93 Buckling-restrained knee braces scheme.

FIGURE 4.94 Cross section of BRBs. Adapted from Xie, Q., 2005. State of the art of buckling-restrained braces in Asia. J. Construct. Steel Res. 61, 727
748.

plate restrained by a square tube, (H) cruciform plate restrained by a square tube, (I) H-shape brace restrained by a square tube, (J) steel plate enclosed by bolted channels and plate, (K) double-tee brace encased by four-connected square tubes, and (L) double-tee double-tube BRB.

4.1 Passive Damping Systems

The encasing tube is one of the major elements affecting the stability of whole system. Takeuchi et al. (2010) studied different tube restrainers through numerical and experimental analyses to investigate the influence of the restrainer tube. Local buckling was studied as a function of width/diameter-to-thickness ratio. The results determined that when the ratio is .65, local buckling is possible, while when the ratio is ,25, no evidence of local buckling failure was found. Moreover, Ju et al. (2009), while studying the influence of tube thickness, found that the compressive force significantly increases (34%
54% greater) for thicknesses of 4
5 mm, compared to a 3 mm one. The other critical material component that affects BRB performance is the filler material. RC or mortar can be utilized. Gheidi et al. (2009) compared the behavior of normal concrete, aggregates, and lean concrete. Test results show that normal concrete, with 25
30 MPa strength, provides the best hysteretic behavior. Filler materials can also be excluded but Gheidi et al. (2009) state that it is safer to use filler material. For debonding agents, several materials have been used in different research campaigns. For example, Wakabayashi et al. (1973) utilized epoxy resin, silicon resin, vinyl tapes, and a composite coating of silicon resin layers on top of epoxy resin. In contrast, Xie (2005) uses silicon painting, styrol foam, polyethene film sheets, and silicone rubber sheets. Given the great amount of possible materials that have been utilized, no clear guideline on the best performing one is available. Besides the brace, other important parameters to take into consideration when designing BRBs are connections and gusset plates. Kersting et al. (2015) provide some recommendations on the type of connections available: connection to gusset plate, connection of the beam to column (including gusset plates), and connection of the column to the base plate. Wigle and Fahnestock (2010) give some recommendations for gusset plates that should be thick enough to prevent large distributed stresses and stiff enough to prevent out-of-plane buckling due to stress concentration. They suggest to have compact gussets since large gussets require thicker welds and are less stiff. Additional recommendations are given by AISC 341 (AISC, 2016a) that require to design gusset plates for 1.1 times the adjusted brace strength in compression since no hinge should develop in the gusset plate. For the end connection, Hussain et al. (2006) state that there are three common layouts, developed by two different BRB manufactures (Fig. 4.95 and Table 4.4): 1. Nippon Steel: standard bolted connection 2. CoreBrace (including Star Seismic): modified bolted connection and true pinned connection In the case of double-tee and double-tube BRB systems, a similar connection of the same type of brace can be utilized (Fig. 4.96; Xie, 2005). In all cases, no splices are accepted inside the BRBs, in order to guarantee reliable plastic behavior of the core. Using BRBs requires testing, following a predefined, detailed loading procedure, described in the appropriate code (e.g., AISC 341 (ASCE, 2016a)).

191

192

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.95 BRB connection type. Adapted from Hussain, S., Benschoten, P.V., Satari, M.A., Lin, S., 2006. Buckling Restrained Braced Frames (BRBF) structures: analysis, design and approvals issues. In: Proceedings of the 75th SEAOC Annual Convention. Long Beach: Coffman Engineering.

Tube-in-tube damper (TTD). The TTD system was proposed by Benavent-Climent (2010) and it consists of two concentric commercially available tubes (Fig. 4.97). The outer hollow section has a series of strips created, cutting a series of slits through the wall and it is connected to the inner slit through a plug-weld. The strips behave as a series of fixed-ended beams and deform in double curvature. The position of the slit is not fixed but can slide along the length of the brace and they can be positioned to facilitate the inspection of the damper.

4.1.1.2.6 Type of devices and manufacturers In the market, there are several manufacturers of displacement-dependent dampers. Structural engineers, architects, or constructors should contact the

4.1 Passive Damping Systems

Table 4.4 BRB Connection Type: Advantages and Disadvantages (Wigle and Fahnestock, 2010) Type of Connection

Advantages

Disadvantages

Standard bolted

• Bigger holes allow more erection tolerance compared to pinned connection. • Multiple bolts provides more redundancy. • Better distribution of forces to the gusset plate.

Modified bolted

• Same as standard bolted. • No splices and fewer bolts. • Longer BRB core length resulting in smaller strains for a given load. • Eliminates overturning moment. • Less installation cost.

• As it is not a true pinned connection, secondary moment and overturning moment form between the connection and brace. • High installation cost due to the number of bolts and splices. • Larger gusset plates to accommodate the bolts, and hence shorter core length, resulting in larger strains compared to pinned connections. • Same as standard bolted.

True pin

• Smaller erection tolerance.

FIGURE 4.96 BRB connection for double-tee and double-tube. Adapted from Xie, Q., 2005. State of the art of buckling-restrained braces in Asia. J. Construct. Steel Res. 61, 727
748.

193

194

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.97 TTD typical damper configuration and equivalent strip geometry. Adapted from Benavent-Climent, A., 2010. A brace-type seismic damper based on yielding the walls of hollow structural sections. Eng. Struct. 32 (4), 1113
1122.

manufactures for the mechanical design of the damping devices. Moreover, most of the standard devices (e.g., ADAS, TADAS, etc.) can be produced by standard steel manufacturers. Selection of the major manufactures and the types of dampers they provide is summarized in the following: • • • •

•

FPC (Italy) produces E-PAD, a special type of BRB. Cast Connex (Canada) developed the cast Scorpion-yielding system that are replaceable modular connectors with stable hysteretic behavior. Corebrace (United States) produces BRBs with different connection types: bolted lug, welded, and pined. Damptech (Denmark) produces rotational friction dampers for earthquake protection and vibration control. Both single- and multilevel dampers are manufactured. Single-level dampers have a constant capacity of 5
5000 kN. Instead, multilevel dampers are suitable for a wider range of vibration control, suitable for both lower vibration amplitudes, for example, induced by wind action or moderate earthquakes and for second level of severe vibrations caused by strong earthquakes. A multilevel dampers are manufactured. Multilevel dampers are suitable for both. FIP Industriale (Italy) produces buckling-restrained axial damper, a special BRB in which the dissipating function is separated by the bracing function, and this is mainly used for the building industry. Moreover, crescent moon,

4.1 Passive Damping Systems

• • • • •

tapered pin (single or double), and U-shaped dampers are mainly used in the bridge industry, as components of seismic isolation systems. FIP Industriale produces SMA devices as well; the first world’s application of SMA for seismic protection of a structure was made by FIP Industriale in the Basilica of San Francesco of Assisi, Italy, during its post-earthquake restoration. Kawakin Core-Tech Co., Ltd. (Japan) produces rotational friction dampers that use a friction pad system. Nippon Steel (Japan) produces unbonded braces and U-shaped steel dampers. Pall Dynamics (United States) develops friction devices. Robinson Seismic Limited (New Zealand) produces lead dampers called penguin vibration damper developed by Penguin Engineering Ltd. Taylor Devices (United States) provides displacement damping devices that produce a specified damping output as a function of position or amplitude. Various configurations are designed based on project-specific requirements.

4.1.2 MASS DAMPING APPROACHES Mass damping approaches consist of applying a dynamic modification system only in few locations in a structure (usually at the top of the structure). Devices can be placed only in one (e.g., TMD) or multiple locations (e.g., MTMD) (Clark, 1988) but without being allocated in a distributed manner, as described for the dynamic modification system in the previous sections. Different devices are part of this category as shown in Table 4.2. The major advantages and disadvantages of each category are summarized in Table 4.5. In the following sections, the major properties and design principles of each device are reviewed in detail (Fig. 4.98).

4.1.2.1 Tuned-mass dampers TMD is a system composed of a mass, spring, and damper (properly tuned) that is attached to a structure to reduce its dynamic response. The original concept was proposed by Frahm (1911) for the ship industry. The main design challenge of this device is to tune its intrinsic frequency to a particular building frequency (usually the fundamental one). Therefore, when the structure is excited with that frequency, the TMD will resonate out of phase with the building and the energy will be dissipated by the damper. The main scope of these dampers is to reduce the wind vibration, but they have also been adopted for seismic protection (Wirsching and Yao, 1973). However, in this case, there are some limitations, such as when the ground motion spectral content is not close to the TMD frequency and when structures behave inelastically (Chey, 2000). Compared to other control devices, TMD involves a relative large mass and displacements. As a consequence, supporting elements are one of the most critical elements in the design process. A TMD is characterized by the following ratios:

195

196

CHAPTER 4 An introduction to dynamic modification devices

Table 4.5 Advantages and Disadvantages of TMD Devices

Advantages

Disadvantages

TMD

• Response to small levels of excitation • Properties can be adjusted in the field • Low maintenance • Cost-effective • Can be designed to add damping to two orthogonal modes of vibration • Effective across all typical tall building periods • Control higher building accelerations than TLDs • Response to small levels of excitation • Mass can be utilized as water supply/storage • Can be designed to add damping to two orthogonal modes of vibration

• Large mass and large space required for installation, but smaller than TLDs of equivalent performance • Effectiveness depending on the maximum mass that can be utilized • Effectiveness depending on the tuning accuracy • Mass, no other functional use

TLD/ TLCD

•

Tuning frequency ratio, f d , as the ratio between the fundamental frequency of the TMD, ωd , to that of the structure, ω. fd5

•

ωd ω

(4.22)

Mass ratio, μ, as the ratio between the mass of the TMD, md , to that of the structure, m. μ5

•

• Damping depending on the screens provided • Water can freeze at low temperature • TLD typically suffers a change in active mass upon tuning • Performance in periods beyond 8 s and/or controlling very high accelerations can be challenging • Possible leakage

md m

(4.23)

Damping ratio, ζ d

The design goal is to find the optimum TMD parameters f d and ζ d for a given μ to reduce the building response (e.g., displacement) under different excitation. The mass ratio is equal to the stiffness ratio (damper stiffness over structure stiffness) in the case that the TMD is in resonance with the structure. For this reason, a proper selection of the mass ratio is the most important aspect. As a general rule of thumb, the mass of a TMD is typically around 0.25%
1.0% of the fundamental modal mass of the building (Chey, 2007), but it can go as high as 5% for

4.1 Passive Damping Systems

FIGURE 4.98 Mass damping approaches simplified schemes.

buildings with high demand damping for serviceability criteria. This relative large mass might require different TMD configurations, depending on the available space (as deeply discussed in the following sections). The concept behind this system is reviewed in Appendix A.5, with the relative TMD equations of motion and their optimized solutions.

4.1.2.1.1 Examples of TMDs In this section, the major applications of TMDs are briefly reviewed, with reference to the major typologies utilized, as follows: • • •

Translational TMDs (vertical and horizontal) Pendulum TMDs (horizontal) Other applications

Translation tuned-mass dampers (vertical and horizontal). In this type of TMD, the mass is supported on rollers (usually bearings), which allow it to translate, and it is attached to the building’s vertical supports with springs and dampers that transmit the out-of-phase motion to the structural vertical members (Fig. 4.99). In

197

198

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.99 Schematic design of a translational TMD.

most of the cases, springs and dampers are provided in both main principal directions, to provide a motion control in two axes. There are several built applications for this type of damper, such as: •

•

•

John Hancock Center (now 200 Clarendon), Boston, has two TMDs applied at the 58th story, consisting of a lead-filled steel box with masses of 2700 kN each. Servo-hydraulic cylinders were utilized to control the motion, and are activated when the horizontal acceleration exceeds 0.003 g for two consecutive cycles. See Section 8.2.2 (Chapter 8) for more details about the case study. Citicorp Center (also known as 601 Lexington), New York City, is equipped with a TMD at the 63rd floor providing an effective damping of 4% with a mass equal to 2% of the effective modal mass of the building, in the first mode of response. The mass is supported on pressure-balance bearings and it activates when the horizontal accelerations exceed 0.003 g for two consecutive cycles. See Section 8.2.1 (Chapter 8) for more details about the case study. Chiba Port Tower, Japan, has TMDs that have a mass of 1/120 compared with the x-direction effective mass, and 1/80 compared with the y-direction. The design damping ratio is 15%.

Pendulum tuned-mass dampers. This TMD solution was introduced to overcome bearing problems, so it supports the mass with cables (Connor, 2003), which makes the system behave like a pendulum (Fig. 4.100). The motion induced by the floor to the pendulum produces an opposite horizontal force. The equation of motion for this system can be expressed as follows (Connor, 2003): FTP sin θ 1 md ðu€ 1 u€ d Þ 5 0

(4.24)

where FTP is the tension in the cable and md is the weight of the pendulum. An additional solution is the rocker pendulum (Connor, 2003), where the restoring force of the cables is generated from the rolling curve surface of the mass. For small θ, the equation of motion is the same as the normal pendulum, with the length of the cable substituted by the surfaces radius.

4.1 Passive Damping Systems

FIGURE 4.100 Schematic design of a pendulum TMD: (A) actual system and (B) equivalent system.

FIGURE 4.101 Park Tower TMD pendulum layout. Adapted from Irwin, P.A., Breukelman, B., 2001. Recent applications of damping systems for wind response. In: Conference Proceedings of the 6th CTBUH World Congress, Melbourne, Australia.

There are several built applications for this type of damper, including the following: • •

•

Taipei 101, Taiwan, has a sphere of 726 tons suspended by eight steel cables looped. See Section 8.2.3 (Chapter 8) for more details about the case study. Shanghai Tower, China, has a simple pendulum with a mass of 1000 tons suspended by steel cables. See Section 8.2.14 (Chapter 8) for more details about the case study. Park Tower, Chicago, is equipped with a pendulum TMD in the tower roof. The mass was approximately 1.4% of the building mass with cable length of 10.35 m. The mass is supported by a tuning frame that can be adjusted in height to control the pendulum period (Fig. 4.101). See Section 8.2.6 (Chapter 8) for more details about the case study.

199

200

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.102 Crystal Tower TMD pendulum layout. Adapted from Connor, J.J., 2003. Introduction to Structural Motion Control. MIT-Prentice Hall Series on Civil, Environmental, and System Engineering.

•

•

Bloomber Tower, New York City, has an inverted pendulum with a dual mass system (with an upper 200- and lower 345-ton steel block). See Ssection 8.2.8 (Chapter 8) for more details about the case study. Crystal Tower, Japan, has a pendulum TMD and it utilizes six cooling and heating storage tanks (Connor, 2003). Each of them has an independent pendulum action and the motion is dissipated with oil dampers attached to them (Fig. 4.102).

Other applications. Particular application of TMD (Fig. 4.103) has been applied in the past such as TLMD with a TMD in one direction and a TLD in the other (Min et al., 2014).

4.1.2.2 Tuned liquid dampers Tuned liquid dampers (TLDs) or tuned sloshing dampers operate in a manner similar to TMD. TLD typically consists of a rigid tank partially filled with liquid (usually water), often rectangular or circular. The liquid contained in a TLD acts as the secondary mass, while gravity provides the restoring (spring) force. The TLD is tuned to slosh at a period nearly equal to the natural period of the structure in order to interact with the structure and reduce its resonant response. Most of the energy dissipated by sloshing dampers is through the wave action and it can be increased with the inclusions of lattice screens and wall roughness (Fig. 4.104). The sloshing dampers can be installed in the form of large tanks (usually from 1 to 4) or in a number of small modules which can be installed in the form of a series of stacks. This makes sloshing dampers significantly adaptable to constricted irregular spaces provided for that purpose. The simple, low maintenance requirements, near-zero

4.1 Passive Damping Systems

FIGURE 4.103 TMD and TLD dual damper experimental setup: (A) TMD and (B) TLCD views (Min et al., 2014).

trigger level, ease of tuning, ability to be designed to operate as bidirectional, and potential dual functionality (i.e., used as a water storage tank) have made the TLD an increasingly popular vibration control device. A challenge in the design of TLD is its nonlinear response. TLD can be classified as either shallow-water, intermediate-water, or deep-water, based on the ratio of the liquid depth (dTL ) to tank length (LTL ) (Fig. 4.104; Tait, 2004). A large portion of the liquid participates in the sloshing motion in a shallow-water TLD (dTL =LTL less than approximately 0.1) and it utilizes wave-breaking as the primary mechanism for energy dissipation. Typically, a large number of tanks are needed in order to achieve the required mass, and the response of a shallow-water TLD

201

202

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.104 Schematic of a rectangular TLD.

becomes increasingly nonlinear as dTL =LTL decreases (Sun et al., 1995). A deepwater TLD (dTL =LTL greater than approximately 0.5) responds in a more linear manner, however less of the contained liquid participates in the sloshing motion, reducing its effectiveness. In addition, dissipation devices, such as screens, poles, baffles (Warnitchai and Pinkaew, 1998), are needed in order to dissipate a sufficient amount of energy. An intermediate-water depth TLD (between 0.1 and 0.5) responds in a more linear manner than a shallow-water TLD, and a greater portion of the fluid participates in the sloshing motion compared to a deep-water TLD. Location of TLDs can be essential in order to counteract torsional response. In this case, it is best to have a pair of tanks as far away from the center of the tower’s center of rigidity as possible. Moreover, it is best to locate the dampers to have their centroid located near the center of the building plan. To ensure that the damper performed to its full capacity, it is highly recommended that onsite measurements are made for the natural frequencies of the relevant modes of vibration prior to commencement of fabrication of the TLD and also immediately prior to installation. It is also prudent to set up long-term monitoring of the relevant natural frequency of the subject building.

4.1.2.2.1 Analytical and numerical fluid models Numerous fluid models have been developed to simulate the response of a TLD. Linear fluid models can be used; however, they are often limited to small liquid sloshing response amplitudes. Nonlinear fluid models can also be employed to capture the nonlinear behavior arising from the free surface boundary conditions. For example, the shallow-water wave theory (Shimizu and Hayama, 1987; Lepelletier and Raichlen, 1988; Sun et al., 1992; Kaneko and Ishikawa, 1999; Tait et al., 2005)

4.1 Passive Damping Systems

and potential flow theory (Frandsen, 2005; Love and Tait, 2010, 2013; Faltinsen et al., 2011) have been successfully used to model a TLD. In addition, several of these models are capable of including the influence of additional dissipation devices, such as screens. The finite-element method, finite difference method, and finite volume method have all been employed to numerically simulate the response of a TLD (Faltinsen and Timoka 2009; Ibrahim, 2005). Smoothed particle hydrodymanics, which is a gridless numerical method, is increasingly being used to model sloshing liquid and has been utilized to model TLDs (Marsh et al., 2010). Although fluid models can capture the nonlinear behavior of the sloshing liquid, they can be challenging to develop and implement.

4.1.2.2.2 Equivalent mechanical models The preliminary design of a TLD is often completed using an equivalent mechanical model. The equivalent mechanical model enables the TLD to be modeled as an equivalent TMD. The sloshing liquid inside simple tank geometries with small response amplitudes can be modeled using a linear spring and mass system (Graham and Rodriguez, 1952; Kareem and Sun, 1987). A dashpot or other energydissipating devices can be added to account for energy dissipation due to viscous losses, wave-breaking, and the presences of screens, poles, etc. (Sun et al., 1995; Warnitchai and Pinkaew, 1998; Tait et al., 2005; Yu et al., 1999; Tait, 2008; Love and Tait, 2011). The mass and spring values for an equivalent mechanical model, which can be either linear or nonlinear, are selected by matching either the energy of the system or resulting forces from the sloshing liquid. The energy of the system or the resulting forces of the sloshing liquid can be evaluated using analytical models (Warnitchai and Pinkaew, 1998; Kareem and Sun, 1987; Tait, 2008; Love and Tait, 2011; Deng and Tait, 2009) or experimental test results (Sun et al., 1995; Yu et al., 1999; Tait et al., 2004). Similar to TMD (as shown in Section 4.1.2.1), the response of a TLD system can be described with the following three parameters: •

Mass ratio 2

μ5

•

φ md m

(4.25)

where φ is the normalized deflection mode shape at the level of the absorber. Tuning ratio, f d , similar to Eq. (4.22) for TMD. The fundamental sloshing frequency of the TLD can be estimated based on linear theory (and small amplitudes) by Lamb (1932), as follows: 1 fd 5 2π

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   πg πdTL tanh LTL dTL

The above relationship is considered to be valid for small sloshing amplitudes and in particular when screens are utilized (Tait, 2004).

(4.26)

203

204

CHAPTER 4 An introduction to dynamic modification devices

•

Damping ratio ζd 5

rffiffiffiffiffiffi 1 ηs ρF Ξ ση 2π md

(4.27)

where ηs is the screen loss coefficient, ρF is the fluid density, ση is root-meansquare free surface motion at the tank wall, and Ξ are parameter related to the tank geometry and the screen configuration, respectively (Deng and Tait, 2009). Additional details about the equivalent mechanical model is shown in Appendix A.6.

4.1.2.2.3 Tuned liquid column dampers Tuned liquid column damper (TLCD) is a special version of TLD with a U-shape, in which the liquid motion in the column counteracts the forces in the structure. Most of the energy dissipated is through viscous dissipation generated by turbulent eddies at the orifice plates and this dissipation can be increased with the inclusions of lattice screens and wall roughness (Fig. 4.105). When resisting motion along one of the translational modes of vibration, it is best to locate the dampers centroid near the center of the building plan. TLCD is not as adaptable to constrictions in space as the TLD. However, it is much easier to have a TLCD double up as a fire hydrant tank as the design can usually be tweaked to fit within a certain length of space or a certain ceiling clearance and remain as a single tank. In the case of bidirectional excitation, two separate TLCDs in each perpendicular direction can be utilized or a bidirectional U-tube can be utilized (Fig. 4.106). In the past alternative, TLCD solutions, instead of the U-shaped, have been utilized. Among all the most common are:

FIGURE 4.105 Perspective view of a TLCD.

4.1 Passive Damping Systems

FIGURE 4.106 Bidirectional TLCD. Adapted from Rozas, L., Boroschek, R.L., Tamburrino, A., Roja, M., 2016. A bidirectional tuned liquid column damper for reducing the seismic response of buildings. Struct. Control Health Monit. 23, 621
640.

• • • •

V-shaped TLCD developed by Gao et al. (1997) LCVA that is a TLCD with a nonuniform column cross section (Hitchcock et al., 1997; Chang and Hsu, 1999) Circular/torsional TLCDs (Huo and Li, 2011) Multiple TLCDs (MTLCDs) that would induce to utilize smaller TLCDs and allow better control of the building motion control (Sadek et al., 1998; Gao et al., 1999)

Similar to TMDs, in the design of TLCDs, the mass ratio concept can be calculated as follows (Sadek et al., 1998): μ5

ρF ATLCD LTLCD m

(4.28)

where ATLCD is the cross-sectional area of the tube; LTLCD is the liquid column length; and ρF is the liquid density. For structures with huge masses, it is practical to use multiple U-tubes to achieve the desired mass ratio; the cross-sectional area of the individual U-tubes can be obtained by dividing ATLCD by the number of tubes (Sadek et al., 1998). In the case where an MTLCD (Fig. 4.107) is required, the mass ratio can be defined as: ρF ATLCD μ5

Nd P j51

m

LTLCD;j (4.29)

where ATLCD is the cross-sectional area of the tube, assumed identical for each group of TLCDs; and LTLCD;j is the liquid column length associated to the jth group of devices. There may be one or several U-tubes in each group of TLCDs such that they meet the required liquid mass ratio. Sadek et al. (1998)

205

206

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.107 SDOF structure with MTLCD.

demonstrated that five groups are found to be the most desirable number for reducing the response. Similar to TLDs, TLCDs also have an equivalent linear model which can be utilized in lieu of more complex nonlinear models (Den Hartog, 1956). A deep discussion about this is given in Appendix A.7.

4.1.2.3 TMD/TLD/TLCD manufactures In the market, there are several manufacturers of mass dampers. Alternately, contractors can decide to self-build the mass for the TMD or the tank for containing the fluid for the TLD and TLCD. Structural engineers, architects, or constructors should contact the manufactures for the mechanical design of the damping devices. A selection of the major manufactures and the types of dampers they provide is summarized in the following: • • • • •

• • • • •

A&H Costum (Canada) produces TMD from the engineering aspects to manufacturing. Deicon (United States) produces a wide range of TMDs, such as pendulum, adjustable TMD. Fiber Technology Corporation (United States) produces water tanks for TLD. FIP Industriale (Italy) produces a wide range of TMDs, such as translational, pendulum. Gerb (Germany) produces vertical mass dampers with threaded coil or tension sprigs. Leaf springs and pendulum suspensions are used for controlling horizontal and torsional vibrations. Hummingbird Kinetics (United States) produces a multi-TLCD system that can be spread-tuned and distributed throughout a building. Maurer (Germany) produces a wide range of TMDs and they have carried out several application with inverted TMD. MTS System Corporation (United States) produces TMD. TESolution (Korea) produces pendulum TMD. Vicoda (Germany) produces pendulum TMD for a frequency building range of 0.1
0.7 Hz and a mass of 11
50 tons with an eddy current damping method.

4.2 Seismic Isolation

4.2 SEISMIC ISOLATION Isolation is one of the most old and well-known protection systems for buildings constructed in earthquake-prone regions. The principle consists of uncoupling the building motion from the input ground motion through isolating the structure from the ground, thus reducing the energy transmitted to the structure. In addition, energy-dissipation devices can be utilized to further reduce the energy transmitted to the structure, and in particular to reduce the horizontal displacement when the seismicity is very high. In the case the isolation is located underneath the building, it is called base isolation. To simply understand how the system works, the ideal example would be a building supported on rollers without friction: when the ground moves, the building will not move (Fig. 4.108). Rather than theoretical rollers, flexible pads are often utilized (base isolators) in such a way that foundation movements are partially transferred to the building. Therefore, the main concept is to increase the buildings’ flexibility to reduce the motion induced in the building (Fig. 4.108). Based on these simple principles, it is clear that careful examination is needed to choose the most suitable device. A practical base isolation requires sufficient horizontal flexibility, damping, recentering, and resistance to sustain service loads (Kelly, 2001b). As a general rule of thumb, base isolation is considered to be most suitable in the following cases (Ariga et al., 2006; Irikura et al., 2004; Kamae et al., 2004; Takewaki, 2005, 2008a,b; Takewaki and Fujita, 2009): • • •

Soil that does not produce long-period predominant ground motions. Isolation level has a high displacement capacity. Low- to medium-rise building resting in hard soil.

FIGURE 4.108 Base isolation schematic concept.

207

208

CHAPTER 4 An introduction to dynamic modification devices

For tall buildings, base isolation is usually considered unsuitable. In most cases, the design of the isolated building has a period of around 2
3 s, i.e., the nonisolated structure should have a fundamental period of , 1.0 s (Jain and Thakkar, 2004), or even less: the rule of thumb to have a good seismic isolation is to have a ratio of 3 between the period of the isolated building and the period of the fixed-base building. However, in the last decades, studies have been conducted on longer period structures (i.e., taller buildings) to understand the influence of base isolation for this building typology: •

•

Jain and Thakkar (2004) proposed the application of base isolation to tall buildings through the utilization of one or more of the following strategies: (1) superstructure stiffening, (2) additional damping in the superstructure, and (3) increased flexibility of the isolation system. The study was conducted on three buildings that range from 10 to 20 stories, equipped with low-damping laminated rubber bearings isolators with viscous dampers. Becker et al. (2015) have shown that in Japan more than 200 base-isolated tall buildings (between 60 and 180 m) have been successfully constructed. One of the major concerns of utilizing base isolation in tall buildings is the need to accommodate large displacement demands at the isolation level, as well as the large overturning that could cause high compression, as well as tension, at the isolation level. In order to overcome problems with uplift, new devices (restrained friction pendulum systems) have been proposed by Roussis and Constainou (2006). Other studies have been conducted with variable-stiffness hybrid devices (depending on the force level, rigidity under wind loads, and flexibility under higher earthquake loads) that could be more suitable for tall buildings design with base isolation (Komuro et al., 2005). Some of the most important applications of base isolation in tall buildings are:

•

• • •

•

Nakanoshima Festival Tower (Osaka, Japan) is a 200-m (39-story) tall building completed in 2012. It features a mid-story seismic isolation systems (16 square lead rubber bearing (LRB) added with oil damper) on the boundary between the podium levels and the above office tower (Nakagawa et al., 2015). A megastruss was built above the isolation level to transfer the tower loads to the system of columns underneath (Fig. 4.109). Tower (unnamed) with a height of 177.4 m in Japan constructed in 2006 (Komuro et al., 2005). Tokio Skystree east tower (Tokio, Japan) is a 158-m (31-story) building completed in 2012. Shinagawa (Tokyo, Japan) is a 155-m (32-story) building completed in 2015. The isolation layer is located at the base level and it consists of combination or rubber bearings and oil dampers. Shiodome Sumitomo Building is a 120-m (25-story) mid-level isolated building in Tokyo constructed in 2004 (Tsuneki et al., 2004). The isolation

4.2 Seismic Isolation

FIGURE 4.109 Nakanoshima Festival Tower (Osaka): building view (A) and section (B) of the isolation scheme (Nakagawa et al., 2015).

•

•

•

level is provided at the 11th floor (i.e., supporting 14th floor), comprises 41 natural rubber bearings (NRBs), 100 lead dampers, and 14 steel rod dampers. LA City Hall (Los Angeles, United States) with a height of 138 m (32 stories) was built in 1926 and retrofitted with 416 HDR and 90 flat sliding polytetrafluoroethylene (PTFE) bearings in 2001. A supplemental damping system composed of 26 viscous dampers (in each direction) and 6 viscous dampers (in each direction) was added at the plane of isolation and between the 26th and the 27th floor, respectively. Thousand tower (Kawasaki City, Japan) is a 135-m (41-story) tall building in the city that was constructed in 2002. The base isolators are located below the first floor (Kawabata et al., 2001). Yozemi Tower (Tokyo, Japan) is a 134-m (26-story) tall building completed in 2008. The isolation system is composed of 25 Rubber Bearing (RB) and 24 flat sliding PTFE bearings. In addition, there are 12 semiactively controlled oil (hydraulic) damper and 12 passive oil dampers. Moreover, the upper structure is equipped with VE dampers located at the annular space between two bracing (EERI, 2012).

209

210

CHAPTER 4 An introduction to dynamic modification devices

•

•

•

•

Shizimu Corporation (Tokyo, Japan) is a 100-m (22-story) tall building completed in 2012. It combines an RC concrete core with 42 seismic isolators (32 LRBs, 10 NRBs, and 10 oil dampers) in between first and second basement floor. J2 Building of Tokyo Institute of Technology (Tokyo, Japan) is a 91-m (20story) tall building. It has an isolation system of 16 NRBs acting in parallel with 15 yielding and 2 oil dampers. Sendai MT Building (Sendai, Japan) is a 85-m (18-story) tall building completed in 1999. It has a combination of 26 low-damping rubber bearings (LDRBs) and 10 flat sliding PTFE bearing below the first level. Sliding bearings are mainly positioned in the interior columns to have the smallest axial fluctuation (Shinozaki et al., 2004). Nunoa Capital Building (Santiago, Chile) has two identical towers and is a 75 m (29 stories) tall building. The towers have an eccentric RC core with perimeter frame and L-shaped wall corners. A common isolation system is utilized at the underground level with 24 NRBs (16 of which are LRBs) (Lagos et al., 2014) (see more details in Section 8.3.1 (Chapter 8)).

4.2.1 BASE-ISOLATION TYPES The modern application of base isolation system to buildings started in 1970s with the pioneering work carried out at the New Zealand Department of Scientific and Industrial Research by Skinner et al. (1975a,b). Afterward, the majority of developments and applications of seismic base isolation happened in few countries such as Japan, New Zealand, United States, China, and Italy (e.g., Kelly, 1986; Buckle and Mayes, 1990; Taylor et al., 1992; Naeim and Kelly, 1999; Higashino and Okamoto, 2006). Generally speaking, the most common isolators can be divided into two major categories: •

•

Bearing • Natural rubber bearing (NRB) • Low-damping rubber bearing (LDRB) • Lead rubber bearing (LRB) • High-damping rubber bearing (HDRB) Sliding • Flat sliding • Friction pendulum system

In Table 4.6, a summary of the main advantages and disadvantages of these isolation bearings is presented. Among the different isolation types described in this section, the most utilized ones are LRBs, for the bearing type, and friction pendulum, for the sliding type, as given in Table 4.7 (specifically looking at the US market (Taylor, 2012)). In Europe and the Mediterranean area, recently the most common type is becoming the pendulum isolator.

4.2 Seismic Isolation

Table 4.6 Advantages and Disadvantages of Isolation Devices Devices

Advantages

Elastomeric bearings

LDRB • Low to moderate instructure accelerations • Simple to manufacture • Easy to model • Response not strongly sensitive to loading

Disadvantages

• Large displacement • Low damping • P-Δ influence

HDRB • Resistance to service loads • Moderate to high damping

• Strain-dependent stiffness and damping • Scragging-change properties

LRB • Wide choice of stiffness and damping • High-damping level Sliding bearings

• Cyclic change in properties

General • • • •

Low profile Resistance to service loads High-damping levels P-Δ influence

• High in-structure accelerations • Properties are function of pressure and velocity and high initial stiffness • Sticking

Flat sliding bearings • Slide plate separates from pad if uplift loads occur • Simple in concept • No strain hardening • Earthquake and structure independent

• No restoring force

Friction pendulum system bearings • Moderate to high damping • Reduced structural response • Relatively wide damping range

• Fixed vibration period • Uplifted structure with motion

In the following sections, a brief description of each type of aforementioned bearings and their mechanical behavior is presented. Furthermore, the basic theory behind base isolation is briefly described in Appendix A.8.

211

212

CHAPTER 4 An introduction to dynamic modification devices

Table 4.7 Approximate Distribution of Seismic Isolation Types in the United States (Taylor, 2012) Isolation System Class

Isolation System Type

Percentage of Total Projects

Bearing

LRBs HDRBs NRBs Friction pendulum system Other types Total

45% 25% 2% 22% 6% 100%

Sliding Other

FIGURE 4.110 Example of rubber bearing (A) and its force-deformation behavior (B).

4.2.1.1 Bearing systems Elastomeric isolators consist of alternating layers of natural or synthetic rubber and steel plates (Fig. 4.110A). The steel plates are needed, so that the rubber does not bulge and the bearings can sustain the high vertical loads and provide high vertical stiffness. However, they do not constrain the horizontal deformations due to shear in the rubber layers; thus, this kind of bearing is more flexible when subjected to lateral loads compared to vertical loads (Kelly, 2001b). A typical lateral force
displacement behavior of NRB is shown in Fig. 4.110B. Despite the easiness of fabrication, such type of bearing does not provide much dissipation, usually 2%
3% damping, and it will probably move under service loads (Kelly, 2001b). For these reasons, low-damping NRBs are used frequently in conjunction with supplementary dissipating systems. For example, in New Zealand, hysteretic devices were added, such as steel dampers (bending-beam and torsional-beam), lead plugs, and lead-extrusion dampers (Buckle and Mayes, 1990; Buckle and Mayes, 1990; Skinner et al., 1993). Similarly, in Japan, the following devices were mainly added to elastomeric bearings: viscous dampers, leadextrusion dampers, steel dampers, and frictional devices (Skinner et al., 1993;

4.2 Seismic Isolation

FIGURE 4.111 Examples of LRBs (A) and its force-deformation behavior (B).

Naeim and Kelly, 1999). Based on these studies, the common solution to increase dissipation is to utilize LRBs or high-damping rubber instead of NRBs. LRBs (Fig. 4.111A) consist of a lead plug force fitted into an elastomeric bearing. The lead core provides dissipation under design loads. The vertical capacity is stable at all loads given from the steel
rubber layer composition. Under lateral displacements, the hysteretic behavior is a combination between the linear elastic behavior of the bearing part and the elastic-perfectly-plastic behavior of the lead core part (Fig. 4.111B). Therefore, the major portion of damping is expected to be obtained from the lead core (Kelly, 2001b). The first application of LRB was for the four-story William Clayton Building, New Zealand, in 1981 (Skinner et al., 1993), while the major and successful application of such a bearing system was in bridges since 1973 in New Zealand (Skinner et al., 1993). Subsequently, it was used extensively in New Zealand, Japan, and the United States (Naeim and Kelly, 1999). High-damping rubber was developed by the Malaysia Rubber Producers’ Research Association (MRPRA) of the United Kingdom (Derham et al., 1985). This was obtained through adding additives to natural rubber, such as extrafine carbon clock, oil, and resins. This is in order to improve its damping capacity up to about 10%
20% (Naeim and Kelly, 1999; Kelly, 2001b). The damping provided by such kind of bearing consists of both hysteretic and viscous components, where the viscous damping contribution (frequency dependent) is lower (Kelly, 2001b). The layout of such a bearing type is generally similar to that of an NRB (Fig. 4.110A), but its hysteretic behavior is much wider resulting in a higher level of damping (Kelly, 2001b).

4.2.1.2 Sliding systems The second major category of isolators is the sliding system, where a sliding element with a given friction coefficient is placed in between the foundation and the base of the structure, limiting the accelerations and forces (Kelly, 2001b). There are two major categories: flat sliding and friction pendulum bearings (Fig. 4.112).

213

214

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.112 Examples of sliding bearings.

Flat sliding bearings (Fig. 4.112A) have been applied in superstructures of many bridges since 1965 (Skinner et al., 1993), such as: • •

Free-sliding bearings, usually pot bearings, in order to provide rotation capability. PTFE disk, backed by a rubber layer positioned into a slot in a steel plate, which slides against a stainless steel surface. The rubber layer is used to have a uniform bearing stress distribution. This is usually a more economical approach than standard elastomeric bearings (Taylor, 2012).

The friction coefficient in such systems under design earthquakes is in the range of 0.1
0.15 for unlubricated bearings and , 0.02 for lubricated bearings (Tyler, 1977). In general, the hysteretic loop of a flat sliding isolator without a restoring force is like a rectangular shape. As a remedy to this, such sliding can be applied in conjunction with some elements that provide restoring force (e.g., LRB or HDRB) or other types of sliding surfaces such as spherical surface can be used (Kelly, 2001b). Friction pendulum bearings (Fig. 4.112B) can be established based on the movement of the weight of the structure (seated on a low-friction slider) over a spherical sliding surface. This kind of bearing combines the sliding mechanism and a restoring force produced by its geometry. In fact, the combination of the lateral displacement and vertical displacement of the main mass leads to a restoring force. One side of the articulated slider which is in contact with the sliding surface is coated with lowfriction material and the other side with stainless steel. Damping is produced due to the friction existing between the slider and the sliding surface. Moreover, the effective stiffness in such systems is a function of the concave surface radius (Naeim and Kelly, 1999). Friction pendulum bearings can include “single” (Fig. 4.113), “double” (Fig. 4.114), “triple” (Fig. 4.115), or “quintuple” (Fig. 4.116; Lee and Constantinou, 2016) isolators depending on the concave surfaces utilized. Furthermore, as mentioned earlier, it is common to add other dissipation devices in parallel (such as viscous dampers). This hybrid system is commonly used to reduce the base isolation displacement limit (Iskhakov and Ribabok, 2007), especially in buildings located within a few kilometers from an active fault (Taylor, 2012). One example of this approach has been utilized in the Regional Medical Center in San Bernardino, California (Taylor, 2012). It is important to note that there are other base isolation systems available, such as seismic ball bearings and steel springs with viscous dashpots (Taylor,

4.2 Seismic Isolation

FIGURE 4.113 Single pendulum system principles (A) and its force
displacement behavior (B).

FIGURE 4.114 Double pendulum system principles (A) and its force
displacement behavior (B).

2012) and active base isolation systems (Chia-Ming and Spencer, 2010). While these technological systems are very interesting, further description is outside the scope of this guide.

4.2.2 BASE-ISOLATION MANUFACTURES Structural engineers, architects, or constructors should contact the manufactures for the mechanical design of the damping devices. In the market, there are several manufacturers of base isolation systems and a selection of the major manufactures and the types of dampers they provide is summarized in the following: •

Bridgestone (Japan) produces multirubber bearings with an outer cover rubber laminations against ultraviolet rays and ozone exposure.

215

216

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.115 Triple pendulum system principles (A) and its force
displacement behavior (B).

FIGURE 4.116 Quintuple pendulum system principles (A) and its force
displacement behavior (B).

•

• •

Dynamic Isolation System (United States) produces a wide range of bearing and sliding isolators. Bearings can range from 0.3 to 1.5 m in diameter and have a capacity of up to 3629 tons. The sliding isolator consist of a PTFE (Teflon) disk on a stainless steel plate and can range from 0.3 to 1.0 m. FPC (Italy) produces a wide range of isolation devices (Isosism), such as sliding pendulum, LRB, and HDRB. FIP Industriale (Italy) produces a complete range of isolation devices, both sliding (both flat with dissipation devices and pendulum isolators) and elastomeric bearings (LDRB, HDRB, LRB). A particular type of sliding isolator combines in the same device a flat slider and steel hysteretic and/or viscous dampers, and has been often used for seismic isolation of bridges. The

4.3 Active, Semiactive, and Hybrid Systems

•

•

•

•

pendulum isolators (curved surface sliders according to European Standard EN 15129 (CEN, 2009)) can be realized with a wide range of friction coefficient. Maurer (Germany) produces a great range of base isolation devices, such as elastomeric, lead rubber (both LDRB and HDRB), sliding pendulum bearings (with proprietary sliding material MSM). Oiles (Japan) produces RB and LRB multilayer natural rubber with an supporting load ranges from 100 to 2000 tons unit with a circular or square design; moreover, flat sliding bearings with Oiles sliding material on the multilayer natural rubber. In addition, Oiles sliding material is used for producing friction pendulum devices. Tensa International (Italy) produces a wide range of HDRB and LRB. There are three compounds of HDRB (soft, normal, and hard) with different levels of shear modulus and viscous damping. Their capacity range from 1000 to 30,000 kN per unit. Vicoda (Germany) produces all the major typologies of base isolation devices, such as HDRB, LRB, and pendulum.

4.3 ACTIVE, SEMIACTIVE, AND HYBRID SYSTEMS The first development of active control systems in civil engineering dates back 40 years ago (Zuk, 1968; Yao, 1972) and significant development has been made in this field since then (Casciati et al., 2012; Fisco and Adeli, 2011a,b). The advantage of active systems, compared to passive ones, is the possibility to vary their properties based on the structure and the external excitation; therefore, they can accommodate variations and uncertainties in the design and loads. In light of this, structures equipped with such devices are called adaptive/ smart structures (Fisco and Adeli, 2011a). In contrast, they require mechanical energy input, which results in an increase in the hardware costs and reliability issues. The development of these active systems is driven by a combination of sensors, controllers, and control actuators (Fig. 4.117): • • •

The sensor measures the displacement along the degree of freedom. The controller determines the appropriate response to be applied. The actuator applies the required force.

At the base of the control laws, there are several aspects that need to be investigated (Casciati et al., 2012), including models, architectural feedback, performance objectives, control methodology, stability, etc. Revising these aspects is outside the scope of this book and interested readers should review the several publications on the subject (see Casciati et al., 2012; Fisco and Adeli, 2011a,b; Datta, 2003; Housner et al., 1997 for an extensive review on the state of art on the subject).

217

218

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.117 Structure with active control. Adapted from Soong, T.T., Spencer, B.F., 2000. Active, semi-active and hybrid control of structures. In: Proceedings of the 12th World Conference in Earthquake Engineering, Auckland, New Zealand.

Depending on the utilization of the measured information, the most common control algorithms are (Soong and Spencer, 2000; Datta, 2003): (1) feedback control or closed-loop control system (sensors on structural response), (2) feedforward control (sensors on excitation), and (3) feedback and feedforward control (sensors on both structural response and excitation). In general, all the active control logics for vibrations suppression are based on several steps, summarized as: • • • •

Identification of the system vibratory state by means of modal filters or observers Definition of the control law that, starting from the vibration level, returns the damping force Actuation of the control forces through a suitable actuator system Evaluation of possible undesired effects associated with the implementation of the logic on a real system

The general scheme of a control algorithm for TMD is shown in Fig. 4.118. Depending on the mathematical equations and operations utilized, the control system can be subdivided as (Datta, 2003) (1) linear, (2) nonlinear, (3) timevarying, (4) discrete time, (5) lumped parameter, (6) distributed parameter, (7) deterministic, and (8) stochastic. Most of the control theories have deterministic control system with lumped parameters and a time-varying control system. All the control methods can be defined with the typical linear equation of motion, as follows (Soong and Spencer, 2000): M U€ ðtÞ 1 CU_ ðtÞ 1 KU ðtÞ 5 J 1 F ðtÞ 1 J 2 FðtÞ

(4.30)

4.3 Active, Semiactive, and Hybrid Systems

FIGURE 4.118 Control algorithm.

where M, C, and K are the N 3 N mass, damping, and stiffness matrices, respectively; U ðtÞ is the N-dimensional displacement vector; FðtÞ is the m-dimensional external load vector; and F ðtÞ is the r-dimensional control force vector. Herein, J1 (N 3 r) and J2 (N 3 m) are the location matrices of the action of the control force and external load vectors, respectively. The control force vector can be seen as a function of the displacement and velocity vector and as a function of the load vector, in the case of a feedbackfeedforward configuration, shown here (Soong and Spencer, 2000): F ðtÞ 5 G1 U ðtÞ 1 G2 U_ ðtÞ 1 G3 FðtÞ

(4.31)

in which the matrices G1 , G2 , and G3 are the control gains, which could be time dependent (Soong and Spencer, 2000). Substituting Eq. (4.31) with Eq. (4.30) leads to the following formula: M U€ ðtÞ 1 ðC 2 J 1 G2 ÞU_ ðtÞ 1 ðK 2 J 1 G1 ÞU ðtÞ 5 ðJ1 1 J 2 G3 ÞFðtÞ

(4.32)

The equation shows how the feedback system controls the structural properties (stiffness and damping) and the feedforward controls of the external forces on the structural system. Among the active control systems, there are two additional classes: hybrid and semiactive control systems. Hybrid control systems are a combination of a passive

219

220

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.119 Structure with hybrid control. Adapted from Soong, T.T., Spencer, B.F., 2000. Active, semi-active and hybrid control of structures. In: Proceedings of the 12th World Conference in Earthquake Engineering, Auckland, New Zealand.

and active control systems (Fig. 4.119), in which a portion of the control is given by the passive system. In this way, the required mechanical energy is lower than an active system. Instead, semiactive control systems require that the control actuators do not add any mechanical energy directly to the structure (Fig. 4.120). Semiactive control systems are sometimes referred to as “smart” systems since they use the advantages of both passive and active systems. This control process is similar to an active system but, instead of applying a force directly to the structure, it allows to control the properties through a passive energy device (Dyke et al., 1996a,b). These systems have an advantage in that they still provide some sort of control during a power failure because they can run on battery power. In the following sections, the major utilized typologies of these control systems are reviewed. The major systems developed up to date are the following (Fisco and Adeli, 2011a,b; Soong and Spencer, 2000): •

•

•

Active systems • Active TMD (ATMD) system (Fig. 4.121) • Distributed actuator Hybrid systems • Hybrid mass damper (HMD) system (Fig. 4.121) • Semiactive control of base isolation systems • Actuators and passive dampers • Semiactive TLCD with passive dampers Semiactive systems • Semiactive controllable fluid dampers • Semiactive stiffness dampers

4.3 Active, Semiactive, and Hybrid Systems

FIGURE 4.120 Structure with semiactive control. Adapted from Soong, T.T., Spencer, B.F., 2000. Active, semi-active and hybrid control of structures. In: Proceedings of the 12th World Conference in Earthquake Engineering, Auckland, New Zealand.

FIGURE 4.121 ATMD systems.

•

• Semiactive TMD systems (Fig. 4.121) • Semiactive TLCD • Friction control devices • Fluid viscous devices Adaptive systems • Adaptive TMD systems (Fig. 4.121)

221

222

CHAPTER 4 An introduction to dynamic modification devices

Table 4.8 Advantages and Disadvantages of Active, Semiactive, and Hybrid Damping Systems Devices

Advantages

Disadvantages

Active

• Properties adjustable based on the excitation input and on the building properties • Smaller dimension compared to similar TMD • Possibility to suppressed multiple mode of vibration • Lower mechanical energy input than an active system

• Usually higher cost compared to similar passive devices • Reliability issues since they need external energy to be effective • Possibility to use them for high level of excitation

Hybrid

Semiactive

Adaptive

• Lower mechanical energy input than an active system, that is added to the passive damper not directly to the structure • Working partially also during power failure (for what regards the passive devices) • Adjustable stiffness and damping in a time frame longer than the actual time period to which the TMD is tuned

• Reliability issues since they need external energy to be effective • Possibility to use them for high level of excitation • Reliability issues since they need external energy to be effective • Possibility to use them for high level of excitation

• Reliability issues since they need external energy to be effective

As briefly stated in this introduction, there are several advantages and disadvantages of active, semiactive, hybrid, and adaptive systems (Table 4.8). These properties will be further reviewed in the following sections.

4.3.1 ACTIVE SYSTEMS In the following section, major active systems are briefly reviewed. Interested readers should refer to the relative reference.

4.3.1.1 Active tuned-mass damper The introduction of the ATMD came from the shortcomings of the traditional TMD. The structural and fundamental frequency cannot be accurately estimated and it also changes during strong motion events. TMDs are more effective when there is a distinct fundamental period of vibration (i.e., not for irregular structures with poorly separated modes). An ATMD consists of an actuator placed between the structure and the TMD, which identifies when a force needs to be applied, computed in real

4.3 Active, Semiactive, and Hybrid Systems

time. The majority of the studies conducted focus on a single ATMD, but multiple ATMDs can be utilized in one structure (Frisco and Adeli, 2011a). For active mass drivers, the counteracting force to increase structural damping is created by an external energy input. Actuators, such as hydro-pulse cylinders or linear motors, accelerate a reaction mass, producing the required control force. Since it is possible to accelerate the mass to a higher level compared to the restoring force of a passive system, a similar reduction for a purely passive system can be achieved with a much smaller reaction mass. This allows a more compact design and, considering that the AMD system does not have to be tuned to a certain lowfrequency, mounting space defining parameters, such as effective pendulum length, do not have to be considered. Furthermore, complex designs to save the overall height of the passive system are not required. The smaller dimension of an AMD system, compared to that of a TMD, is the main advantage for the application in tall buildings. Another advantage is the possibility to address multiple modes, which become relevant for supertall buildings that display natural frequencies for the second lateral bending modes (Modes 3 and 4) within a critical frequency range.

4.3.1.2 Actuators Actuators are devices that can produce a given force or strain. If these are only applied in the longitudinal direction, the actuators are called linear and they can be hydraulic, electric, and electrohydraulic (Adeli and Saleh, 1997). These are usually placed at member ends (Fig. 4.122, Adeli and Saleh, 1997). If one

FIGURE 4.122 Example of position actuator in truss-bridge structure. Adapted from Adeli, H., Saleh, A., 1997. Optimal control of adaptive/smart bridge structures. J. Struct. Eng. 123 (2), 218
226.

223

224

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.123 Active tendon control of cable-stayed bridges. Adapted from Achkire, Y., 1997. Active tendon control of cable-stayed bridges. Thesis. Universite´ Libre de Bruxelles.

actuator is not enough, two can be placed, one at each end of the member. For large structures with a lot of members, it is important to place the actuators in their optimal position and a lot of combinations are possible (Adeli and Saleh, 1997). One of the most interesting applications of this technology is for floor vibration reduction (Hanagan and Murray, 1997). It is also crucial to mention the importance of the actuator saturation, i.e., actuator reaching its force limit. Several authors have proven that there are no detrimental effects to structural stability in these cases (Agrawal et al., 1997). Another category are active tendons systems, composed of prestressed tendons applied in between floors of a structure, and have variable tension controlled by actuators. One recent application has been developed for a cable-stayed bridge (Fig. 4.123; Achkire, 1997). Another interesting application of tendons is where the stiffness of one building help to control the response of another building by coupling them (Christenson et al., 2003).

4.3.2 HYBRID SYSTEMS To increase building performance, another system consists of combining passive, active, and semiactive systems, referred to as a hybrid system. Although this may seem complex, these methods allow the system to take advantage of the unique benefits that passive, active, and semiactive systems have.

4.3 Active, Semiactive, and Hybrid Systems

FIGURE 4.124 HMD example configuration scheme. Adapted from Maebayshi, K., Shiba, K., Mita, A., Inada, Y., 1992. Hybrid mass damper system for response control of building. In: Proc. of Tenth World Conference on Earthquake Engineering, Rotterdam, The Netherlands.

4.3.2.1 Hybrid mass damper A hybrid control system consists primarily of a passive TMD system and additional actuators (Fig. 4.124; Maebayshi et al., 1992). Contrary to an AMD system, the effective masses of hybrid systems are not directly driven by the actuators, but resonance effects are being used to achieve the required control force. The effectiveness of a TMD can be increased by attaching an auxiliary mass and actuator to the tuned mass, and driving the auxiliary mass with the actuator so that its response is out of phase with the response of the tuned mass. The effect of driving the auxiliary mass is to produce an additional force that complements the force generated by the tuned mass and, therefore, increases the equivalent damping of the TMD. It is possible to obtain the same behavior by attaching the actuator directly to the tuned mass, thereby eliminating the need for an auxiliary mass. The practical objective of an HMD is to shorten the reaction time of a passive system or to introduce a sufficient activation force for passive systems subjected to friction. It is important to note that another HMD was proposed by Fisco and Adeli (2011a,b) that is composed of an active mass damper with passive dampers.

4.3.2.2 Semiactive control of base isolation systems Usually base isolation is not suitable for long building periods, as is the case of tall buildings (see Section 4.2 for a detailed description of tall building with base isolation). For this reason, base isolation systems have been integrated with active or semiactive control systems (Fig. 4.125; Gavin et al., 2003).

225

226

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.125 Base-isolated structure with controllable hydraulic device. Adapted from Gavin, H., Alhan, C., Oka, N., 2003. Fault tolerance of semiactive seismic isolation. J. Struct. Eng. 129 (7), 922
932.

4.3.3 SEMIACTIVE SYSTEMS 4.3.3.1 Semiactive tuned-mass dampers Semiactive tuned-mass dampers (SATMDs) are also capable of adjusting their stiffness and damping parameters, but the adaptation process occurs within the vibration period of the system (contrary to the adaptive TMD). The difference to an AMD is the effective mass of the system that is driven by inertia, without an external force input. The motivation to use SATMDs is mainly to enhance the performance of the control device for stochastic excitation (seismic or gust loading) because the response of the effective mass can be adapted to the input. Based on the excitation, the damping of the system can also be adapted to an optimal value and, in the case that the displacement of the system’s effective mass is too large, the damping can also be increased to maintain a certain mass travel distance. The semiactive control is either addressing the tuning frequency of the TMD or the damping device. With the control of the tuning frequency, it can be ensured that the frequency ratio between the dynamic response of the tower and the TMD tuning frequency remains in a very narrow band for which a TMD with an internal damping ratio lower than the optimum value according to the previously mentioned criteria could achieve a better reduction (Fig. 4.126). Alternatively, the damping of the TMD can be controlled to be small when the frequency ratio is determined to be close to 1 (also see Fig. 4.126). Semiactive control devices are also often viewed as controllable passive devices. More specifically, their resistive or dissipative force produced via control

4.3 Active, Semiactive, and Hybrid Systems

FIGURE 4.126 SATMD strategy to increase performance (reduction) for a smaller frequency band than for a passive TMD due to adjustable frequency/damping.

of internal mechanisms based on external sensor feedback. Hence, they can combine the best elements of active and passive systems, or, in contrast, mitigate their less undesirable features. Because of their low power requirements and large controllable force capacity, semiactive systems provide an attractive alternative to active and hybrid control systems. Semiactive control strategies are particularly promising in addressing many of the challenges to wider application of control to civil engineering structures, by offering the reliability of passive devices, while maintaining the versatility and adaptability of fully active systems. Studies have shown that appropriately implemented semiactive damping systems perform significantly better than passive systems and have the potential to achieve, or even surpass, the performance of fully active systems. Examples of semiactive devices include variable orifice fluid dampers, controllable friction devices, variablestiffness devices, and controllable fluid dampers.

4.3.3.2 Semiactive controllable fluid dampers In recent years, semiactive devices—employing orifices of variable openings or adjustable screens—have been considered by the tall building industry (e.g., Yalla et al., 2001), especially within the framework of conventional TLCDs. Unfortunately, the technical challenge to provide efficient and effective valves capable of quickly responding to external changes in the excitation has not been particularly successful. This lack of success is one of the main reasons behind the growing interest in the research of controllable fluids specifically for semiactive applications. Until now, two main types of “smart” fluids have been considered in the control of structural vibrations: electrorheological (ER) fluids and

227

228

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.127 Schematic diagram of a shear mode MR damper. Adapted from Dyke, S.J., Yi, F., Frech, S., Carlson, J.D., 1999. Application of magnetorheological dampers to seismically excited structures. In: International Modal Analysis Conference, IMAC XVII, FL.

magnetorheological (MR) fluids (Jagadish and Jangid, 2009). ER and MR fluids are suspensions of micrometer-scale particles which, when respectively subjected to an electric or magnetic field, are capable of significantly increasing their apparent viscosity. The yield stress of ER and MR fluids—in their active state—can be accurately controlled by varying the intensity of the electric or electromagnetic field, respectively. This allows response time of the order of milliseconds to be achieved while employing rather low-voltage power sources. Furthermore, in the event of power loss, the damper acts as a passive system, so it still provides some protection. Considering the cost of ER or MR dampers, it would be uneconomical to install the semiactive control devices on every floor of the building: determining the optimum location of such devices within the structure is therefore very important and the genetic algorithm is often used as an efficient optimization method (Yoshida and Dyke, 2005). A prototype shear mode MR damper is shown in Fig. 4.127 (Dyke et al., 1999) and a schematic diagram of this device is shown in Fig. 4.128. Fig. 4.129 shows the force
displacement and the force
velocity response hysteresis loops of an MR damper undergoing a 1.5-Hz sinusoid with an amplitude of 1.5 cm at different constant voltage levels. Considering the advantages of “smart” fluids, there is no doubt that semiactive ER and MR dampers can offer a very viable solution to the ongoing challenges posed by structural vibrations.

4.3.3.3 Semiactive stiffness dampers This type of damper consists of a variable valve fluid piston that is controlled by a motor, in such a way that the damping coefficient of the device can be adjusted in real time. A variety of different solutions have been proposed in the past

4.3 Active, Semiactive, and Hybrid Systems

FIGURE 4.128 Mechanical model of the MR damper.

FIGURE 4.129 Typical MR damper responses. Adapted from Yoshida, O., Dyke, S.J., 2005. Response control of full-scale irregular buildings using magnetorheological dampers. J. Struct. Eng. 131 (5), 734
742.

including resetting semiactive stiffness dampers (Jabbari and Bobrow, 2002); switching semiactive stiffness dampers (Agrawal et al., 2003); and variable-slip force semiactive stiffness dampers (Nishitani et al., 2003).

4.3.4 ADAPTIVE TUNED-MASS DAMPER SYSTEMS Adaptive TMD systems are capable of adjusting either stiffness or the damping coefficient of its components, which helps it adapt to changes in the main system. Compared to SATMDs, the change in these parameters happens in a time frame

229

230

CHAPTER 4 An introduction to dynamic modification devices

FIGURE 4.130 Adaptive radius TMD systems. Adapted from Nagarajaiah, S., 2009. Adaptive passive, semiactive, smart tuned mass dampers: identification and control using empirical mode decomposition, Hilbert transform, and short-term Fourier transform. Struct. Control Health Monit. 16 (7–8), 800
841.

that is longer than the actual time period to which the TMD system is tuned for which means that as long as the system can be subsequently adjusted, and does not require a continuous external energy input, it can be considered an adaptive TMD system. Often the stiffness parameter can be adjusted by changing the amount of tuning springs in the semiactive damper continuously and changing the independent variable stiffness (radius of device or the effective pendulum length, Fig. 4.130). The damping coefficient can be adjusted by using variable orifices (VDDs) or by adjusting the shear area of a dashpot damper (Nagarajaiah, 2009).

4.3.5 CONTROL STRATEGIES The primary strategy to determine the successfulness in the implementation of active systems is through the utilization of an effective control algorithm. This permits the control of the magnitude of the forces to be applied to the structure. Some of these algorithms, including the linear quadratic regulator (LQR) and linear quadratic Gaussian (LQG), have been developed in other fields (e.g., aerospace engineering), but recently several of them have been developed specifically for civil structures (Fisco and Adeli, 2011b). The most important that have been utilized in the past include the following: •

LQR: Originally developed for the aerospace industry, the LQR is one of the most used to solve optimal control problems. This is based on minimizing a cost function, referred to as displacement in civil engineering calculations, and the acceptable level for story acceleration. This method has been

4.3 Active, Semiactive, and Hybrid Systems

•

•

•

•

modified through the past decades to satisfy civil engineering needs (Fisco and Adeli, 2011b). LQG: LQG combines a linear quadratic estimation with an LQR and it was originally intended for systems distributed by white Gaussian noise (Fisco and Adeli, 2011b). Neural network controllers: Neural network controllers are more versatile for practical applications, but they are optimal when controlling the structural response (Datta, 2003). Fuzzy logic control: Fuzzy logic control defines logical variables with continuous values, as opposed to discrete values in classic logical problems (Fisco and Adeli, 2011b). Sliding mode controllers: Sliding mode controllers use high-frequency switching to provide control (Fisco and Adeli, 2011b).

A detailed discussion about these control strategies is out of the scope of this book and interested readers should refer to the appropriate sources on the subject.

4.3.6 FUTURE DIRECTIONS Despite advancements in research on control strategies, there are still some issues in their implementation for full-scale application. One of the major concerns is their performance for high levels of excitation (e.g., large earthquake motions), because active control devices currently developed are primarily used for human comfort requirements (e.g., wind and moderate earthquake excitations). The other major issue is related with the economic implications of their development, construction, and maintenance. Moreover, if the feedback control system (e.g., force feedback control or velocity feedback control) is not precisely designed, the system may be subject to instability resulting in unstable structural responses. Therefore, the control gains should be chosen carefully to maintain stability (Inman, 2006). In general, velocity feedback systems are the most inherently stable ones, especially if the velocity sensor and force actuator are collocated, i.e., they are attached at the same point (Fuller et al., 1996). Furthermore, there are other time application problems (Datta, 2003): 1. 2. 3. 4.

Modeling errors (e.g., limited degrees of freedom). Time delay between the sensing and actuator response. Sensor and controller cannot be applied at all points. Parameter uncertainties and system identification (e.g., material strength, nonlinear behavior). 5. The time control is discrete while the algorithm provides a continuous one. The future direction of this field includes the development of active structures, not only active controlled structures. The first one is where the active and passive elements work together in a synergic structural system. The second one relates to

231

232

CHAPTER 4 An introduction to dynamic modification devices

a structural system that is fitted with active control devices that are activated under extraordinary loads (Soong and Spencer, 2000).

4.3.7 ACTIVE, SEMIACTIVE, AND HYBRID DAMPERS MANUFACTURES In the market, there are different manufacturers of active, semiactive, and hybrid dampers. Structural engineers, architects, or constructors should contact the manufactures for the mechanical design of the damping devices. A selection of the major manufactures and the types of dampers they provide is summarized in the following: • •

• •

Maurer (Germany) produces MR fluid semiactive dampers. Deicon (United States) produces active pendulum TMDs with hydraulic actuator in in place of (or in conjunction with) viscous dampers. A supervisory control scheme continuously readjusts the parameters of the active pendulum to make sure they are always tuned. Gerb (Germany) produces active, semiactive, and hybrid TMDs. Tesolution (Korea) produces HMDs.

4.4 COMPARISON OF DAMPERS IN TALL BUILDINGS In the previous sections, the most common passive and active/semiactive damping devices used in the protection of tall buildings have been reviewed. As it can be seen, the designer has a lot of different devices to choose from and it is hard to find which devices is most suitable for a building. For this reason, the major advantages and disadvantages for each device category are summarized, based on what was discussed in depth in the previous sections. However, it is important to state that each building will have unique and different peculiarities and requirements that would lead to a case-specific investigation in order to determine which device is optimal. Therefore, the designer should choose the damping devices based on the specific project features. In Table 4.9, the major features for each damping device are summarized to help designers, architects, contractors, building owners, etc. determine their initial decision for the tall building under investigation. Among all the above summarized major features, for each device category, another important consideration should be related with the region of the world where the building will be constructed, primarily because some devices would be more economical and readily available in one region compared to another. Moreover, as the research would progress, some new devices can be proposed and existing ones will be further developed from both mechanical and design point of views.

4.4 Comparison of Dampers in Tall Buildings

Table 4.9 Technical Comparison of Damping Devices for Tall Buildings

Devices Fluid viscous damper

Lateral Excitation Resistance Wind/ seismic

Major Advantages

Major Drawbacks

Temperature-independent properties Large damping values

Possible fluid leakage

Small size for very high force

VE damper

Wind/ seismic

Metallic damper

Seismic

Friction Damper

Seismic

Tunedmass damper

Tuned liquid damper

Wind (sometimes seismic)

Wind (sometimes seismic)

Instantaneous added damping to the structure at low amplitudes Linear behavior (easy modeling) No maintenance requirements Incorporated into main structural system Stable and well-understood hysteretic behavior Large energy dissipation per cycle Can be reused after seismic event On-site tuning possible Can be used for most structural forms

Easy interface between structural designer and TMD designer Similar advantages as for TMD Low maintenance and can sometimes be utilized as water supply

Some designs have been found to be unstable at low amplitude vibrations No resistance to static load Challenging to incorporate adequate local stiffness and mechanical details to ensure effectiveness for wind events that result in very small relative motions Temperature- and frequency-dependent properties

Damaged after an event and may require replacement Sliding interface conditions may change and no restoring force May require adjustment after earthquake Large space take at top of building Not suitable for damping strong winds or earthquakes without special measures Maintenance required

Low damping just for the fundamental frequency Large space requirements

(Continued)

233

234

CHAPTER 4 An introduction to dynamic modification devices

Table 4.9 Technical Comparison of Damping Devices for Tall Buildings Continued

Devices

Lateral Excitation Resistance

Major Advantages

Major Drawbacks Cycling change of properties Hard to effectively configure in taller buildings Fixed vibration period and uplift concern Hard to effectively configure in taller buildings Reliability Possibility to use them for high level of excitation, although needs reliable power supply

Rubber bearing isolator (LRB)

Seismic

Wide choice of stiffness and damping

Friction bearing isolator (pendulum)

Seismic

Wide damping range

Active

Seismic/ wind

Adjustable properties Can damp multiple modes

Seismic/ wind

Take smaller space than TMD Lower mechanical input energy

Hybrid/ semiactive/ adaptive

Reliability Possibility to use them for high level of excitation

CHAPTER

Design procedures for tall buildings with dynamic modification devices

5

CHAPTER OUTLINE 5.1 Available Codes and Design Tools.....................................................................236 5.1.1 Codes and Guidelines .....................................................................236 5.1.2 Practical Design Aspects .................................................................248 5.1.3 Structural Analyses.........................................................................251 5.2 Passive Damping Systems ................................................................................264 5.2.1 Step-by-Step Procedure for Distributed Dampers...............................265 5.2.2 Step-by-Step Procedure for Mass Dampers .......................................321 5.3 Isolation Systems.............................................................................................347 5.3.1 Step-by-Step Procedure for Base Isolation ........................................347 5.4 Active, Semiactive, and Hybrid Systems ............................................................365 5.4.1 Literature Review ...........................................................................365 5.4.2 Step-by-Step Procedure ..................................................................367 5.5 Retrofit of Existing Buildings.............................................................................400 5.5.1 Code Requirements ........................................................................400 5.5.2 Evaluation Procedures Based on ASCE 41-13 (ASCE, 2013) .............402 5.5.3 Step-by-Step Procedure ..................................................................403 5.5.4 Case Study: Retrofitting Examples of High-Rise Building with Damping Systems...........................................................................422 5.6 Dynamic Modification Devices Strategy Optimization..........................................429 5.6.1 Introduction ...................................................................................429 5.6.2 Algorithm-Based Optimization Procedures ........................................430 5.6.3 Nonalgorithm-Based Optimization Procedures...................................433

This chapter of the book gives a general overview of the current state-of-the-art design procedures for tall building with dynamic modification devices under wind/ seismic loading. Code prescriptive procedures are reviewed, by examining national and international codes/standards available worldwide. US codes are discussed in detail since they are arguably the most advanced standards for the design of dynamic modification devices. Practical design and structural analysis aspects are also discussed based on the structural engineering office experience in the design of tall buildings with dynamic modification systems. Subsequently, detailed stepby-step procedures, based on code requirements, are then outlined for the major devices categories as discussed in Chapter 4. It is important to note that code Damping Technologies for Tall Buildings. DOI: https://doi.org/10.1016/B978-0-12-815963-7.00005-1 Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved.

235

236

CHAPTER 5 Design procedures for tall buildings

prescriptive procedures are not available for all devices, such as mass dampers (i.e., tuned mass dampers (TMDs), tuned liquid dampers (TLDs), tuned liquid column dampers (TLCDs)). In these cases, an extensive literature review is carried out to summarize alternative noncode prescriptive design procedures.

5.1 AVAILABLE CODES AND DESIGN TOOLS In the following sections, the major codes and guidelines available worldwide for the design of structures with dynamic modification devices are reviewed. In most of the cases, these were not written exclusively for tall buildings. For this reason, when reviewing the step-by-step procedures, in addition to code-based prescriptive procedures (where available), relevant literature is reviewed.

5.1.1 CODES AND GUIDELINES Despite the large variety of research and applications of dynamic modification technology for the seismic design, wind design, and rehabilitation of buildings, not many guidelines have been developed. In reality, most guidelines appear to focus on providing simple recommendations for testing procedures of devices, without any specification for their design. The European code (Eurocode 8 (CEN, 2003)), Italian standards (NTC, 2008), Japanese building code (BCJ, 2013; JSSI, 2003, 2005, 2007), and New Zealand standards (NZS, 2006) provide basic recommendations for structures with added damping, but have no specific recommendations for practitioners. In contrast, in the United States, there are guidelines available for both new (ASCE 7-16 (ASCE, 2017a)) and existing buildings (ASCE41-17 (ASCE, 2017b)). For these reasons, and for most devices, the requirements from US guidelines are utilized as the basis for the design of structures with dynamic modification devices, as will be shown later.

5.1.1.1 European code Eurocode treats dynamic modification system only in the seismic portion (CEN, 2003). In particular, in Part 1 (CEN, 2003), the requirements for base-isolated structures as a means to protect buildings against earthquake actions are discussed. Other types of dynamic modification system are not treated even if they are allowed in Part 3 (CEN, 2003) for the seismic rehabilitation and retrofit of buildings. The code allows to use different types of isolation devices, such as elastomeric bearings and friction pendulum bearings, as long as they are capable of providing one of the following functions (or a combination of them): • • • •

Sustaining vertical loads with high vertical and low horizontal stiffness, allowing large horizontal displacements Energy dissipation through hysteretic and/or viscous mechanisms System recentering Restraining lateral movement for service conditions (not seismic)

5.1 Available Codes and Design Tools

In addition, CEN (2003) provides several general design recommendations, such as: • •

• • •

•

Only the isolation system may reach the ultimate capacity while the structure above and below has to remain elastic at all times. The isolation system should be positioned such that its center of rigidity should be in proximity to the center of rigidity of the structure, above the isolation system itself. A rigid diaphragm shall be provided below and above the isolation interface. No part of the isolation system shall be in tension. Increased reliability is required when isolation devices are utilized. For this reason, an amplification factor of the seismic design displacement is considered. Restrictions are also placed on the soil and the differential movements of the surrounding systems.

For the modeling and analysis of structures with isolation systems, the code (CEN, 2003) emphasizes the importance of estimating and modeling the property variation during the life span of the devices, with a maximum variation of 20%, which is considered acceptable. For these reasons, isolators should be modeled with appropriate constitutive properties. However, a simplified linear model can be utilized when the following requirements are satisfied: • • • •

The equivalent stiffness of the system is greater than 50% of the secant stiffness at 20% of the design displacement. The equivalent total viscous damping is less than 30%. The force
displacement characteristics of the isolation system have a variation less than 10% of the design values for a range of around 30%. The increase in the restoring force between 50% and 100% of the design displacement is not less than 2.5% of the total gravity load on the isolation system.

If these requirements are satisfied, the isolation system can be modeled with an equivalent stiffness at the design displacement, for the considered limit states (of the base-isolated building) into account the dissipation energy of the devices. For higher modes, no additional supplemental viscous damping shall be included. In case both stiffness and damping are highly dependent on the design displacement, an iterative procedure should be carried out until a difference of 5% between assumed and calculated damping is reached. The type of analysis permitted for structures with isolation devices depends on the satisfaction of the following criteria: •

Static linear analysis is allowed when the following conditions are satisfied: • Isolation system can be linearly modeled. • Isolated structure equivalent period is in between 3 times the fixed-base building period and 3 seconds.

237

238

CHAPTER 5 Design procedures for tall buildings

•

•

Vertical stiffness is greater than 800 times the horizontal stiffness of the isolation system. • Vertical period is less than 0.1 second. • There is no tension in any of the isolation devices. • The structural system is regular. • Height of the structure above the isolation system is less than 20 m or 5 floors. • Structure below the isolation can be considered rigid, if the period is less than 0.05 second. • Major plan dimension is less than 50 m. • Eccentricity shall be less than 3% in any direction. If all of these requirements are satisfied, two models can be developed: one for the superstructure and one for the substructure. Dynamic linear analysis is allowed when the isolation system can be modeled as a linear system subject to the requirement that follows: • The model shall consider both the substructure and the superstructure unless the substructure is just the foundation system. • Response-spectrum (RSA) or response-history analyses (RHA) can be utilized. For RSA a bidirectional combination shall be utilized and the vertical component is required when the vertical stiffness is less than 800 times the horizontal stiffness of the isolation system. The spectrum can be reduced based on the total damping. For the RHA the total damping shall be defined for each mode, when a modal decomposition approach is utilized, or through modifying the damping matrix.

As can be seen the requirements for utilizing static linear analysis exclude the possibility of applying it to tall buildings. Therefore, for tall buildings, only the dynamic linear analyses or nonlinear analyses (both static and dynamic) can be utilized.

5.1.1.2 Italian code The Italian building code, called NTC (2008) (“Nuove Norme Tecniche”), does not give specific design criteria for dynamic modification systems, but Section 7.10 provides some general guidelines for the design of isolation systems positioned underneath the main building. These guidelines follow the same requirements as Eurocode 8 (CEN, 2003) as reviewed in the previous section.

5.1.1.3 Japanese code The Japanese Building Standard Law (BSL) (BCJ, 2013) provides some general guidelines for added damping systems and relative calculations based on energy balance theory (see Chapters 3 and 4). The main concept is that dissipation devices should absorb enough energy to not to reach the main structural damage limit state. However, the Japan Society of Seismic Isolation (JSSI, 2003, 2005, 2007) proposed simplified procedures for four different damper types (Fig. 4.1 (Kasai et al., 2008)) and nine different frame configurations (Fig. 5.1 (Symans et al., 2008)). The proposed analyses consider the building being modeled as a multiple

5.1 Available Codes and Design Tools

FIGURE 5.1 Possible JSSI frame configurations. Adapted from Kasai, K., Kibayashi, M., 2004. JSSI manual for building passive control technology part-1 manual contents and design/analysis methods. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1 6, 2004 Paper No. 2989.

FIGURE 5.2 MDOF shear beam model for flexural behavior.

degree-of-freedom (MDOF) stick frame system, which allows it to also be utilized if significant flexural behavior is present in the building (Fig. 5.2 (Kasai et al., 2008)). Structures equipped with these dissipative devices can be idealized at each story, according to JSSI (2003, 2005, 2007), with an equivalent single

239

240

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.3 SDOF-added component models.

degree-of-freedom (SDOF) system (Fig. 5.2) consisting of damper, supporting member (e.g., brace), and frame. The damper and relative supporting members are called the “added components” (Kasai and Kibayashi, 2004). Fig. 5.3 shows the model for four different dampers: steel, viscoelastic, and viscous/oil. The models are composed of a spring with stiffness (K), dashpot, and viscous coefficient (C). Furthermore, the subcaption b refers to the brace and d to the damper. All the models have elements in series except the viscoelastic device, where the damper stiffness and viscous parts are in parallel. The elastic stiffnesses of the oil and that of the viscous dampers are due to the compressive modulus of the oil and viscous liquid, respectively. In these cases, the equivalent brace stiffness is given by adding the brace and damper stiffness. The only different case is the steel damper, where a unique stiffness is defined. It is important to note that the Japanese code states the importance of modeling the frequency dependency of the viscoelastic properties. JSSI (Kasai and Kibayashi, 2004) also provides recommendations for the hysteretic characteristics of each device for the peak deformations of the damper, the added component, and the whole system (Fig. 5.4). From the hysteresis loop, it is 0 00 possible to compute the storage (K ) and loss (K ) stiffnesses computed by dividing peak force by the corresponding displacement (black dot) and zero displacement force (white dot) to the peak displacement, respectively. The response and design of structures equipped with these devices have been simplified by JSSI (2003, 2005, 2007) by proposing performance curves (Fig. 5.5 (Kasai and Kibayashi, 2004)). These are a continuous function of an SDOF system based on damper properties, storage, and loss stiffness for an idealized response spectrum. The curves show both the displacement (Rd ) and the force (acceleration) reduction ratio (Ra ), defined as ratio between the peak responses with and without damper. In Fig. 5.5, Kf is the frame stiffness, μ is the ductility 00 of the steel damper, and Kd1 is the damper loss stiffness. The curves show how, up to a certain point, larger dampers lead to more reduction, at which point

FIGURE 5.4 JSSI dampers hysteretic properties for three different peak deformations. Adapted from Kasai, K., Kibayashi, M., 2004. JSSI manual for building passive control technology part-1 manual contents and design/analysis methods. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1
6, 2004 Paper No. 2989.

242

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.5 Performance curves for four different damper types. Adapted from Kasai, K., Kibayashi, M., 2004. JSSI manual for building passive control technology part-1 manual contents and design/analysis methods. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1
6, 2004 Paper No. 2989.

they become no longer beneficial. Moreover, when a brace has low stiffness and large damper forces, the energy dissipation becomes smaller. Based on the performance curves, as shown in Fig. 5.5, dampers can be designed as an SDOF system. This can be extended to multistory systems, in case the story-to-damper stiffness ratio satisfies the SDOF system model. This should be obtained by considering 80% of the total mass of the building and the effective height is based on the first mode deflected shape. Kasai and Kibayashi (2004) show a design step-by-step procedure as shown in Fig. 5.6. It is important to note that design validation should be carried out through time-history analyses with modeling of damper devices (Kasai and Kibayashi, 2004).

5.1 Available Codes and Design Tools

FIGURE 5.6 Damper and system design procedure. Adapted from Kasai, K., Kibayashi, M., 2004. JSSI manual for building passive control technology part-1 manual contents and design/analysis methods. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1
6, 2004 Paper No. 2989.

5.1.1.4 Chinese code The Chinese seismic code, GB 50011 (GB50011, 2010), provides design recommendations for seismic isolation and seismic energy
dissipating systems, in Chapters 3.8 and 12, respectively. Isolation systems are permitted for multistory masonry and reinforced concrete (RC) frame buildings, if the following requirements are satisfied: • • • •

Regular structures can be designed according to provision of Section 5.1.2 of GB 50011 (GB50011, 2010). Irregular structures adopting seismic isolation need to be verified by model tests. Nonseismic lateral action shall not be greater than 10% of the total structural weight. Components of the seismic-isolated story have enough stiffness and damping to avoid interrupting any mechanical equipment.

Energy dissipation systems are permitted for steel and RC structures. These systems must provide sufficient damping and shall be designed per the provisions for each structural type. Moreover, these systems can be utilized for the buildings having special performance requirements or with a seismic intensity between 8 (20:2=0:3 g) and 9 (0.40 g).

243

244

CHAPTER 5 Design procedures for tall buildings

The following is a brief overview of the requirements for both seismicisolation and energy dissipation systems. It is important to note that GB50111 (2010) requires to carry out performance-based design (PBD) for structures equipped with both damper and isolation systems according to the requirements for seismic performance objectives (POs) (as reviewed in Chapter 3).

5.1.1.4.1 Base-isolated buildings The code defines as isolation stories where isolators and related elements are positioned in within the structure. The general requirements for the design of these systems are as follows: •

A shear-type model can be utilized for the structure above the isolation system. Time-history analyses shall be performed with near-fault input accelerations, and in case near-field accelerograms are not available, the analyses shall be performed with standard acceleration with the following multiplication factors: 1.5 when the site is located within 5 km from the fault and 1.25 when the site is located beyond 5 km from the fault. In the case of masonry structures, a simplified procedure can be utilized [Appendix L of GB50011 (2010)].

•

•

The provisions allow the use of rubber isolator units with the following requirements: •

The ultimate displacement, under the compression stress limits given in Table 5.1, shall not exceed the maximum of either 0.55 times the effective diameter or 3 times the total rubber isolator thickness. In case the secondary form factor (ratio of the diameter and the total rubber thickness) of the isolator unit is less than five, the stress limits shall be reduced as follows: by 20% if between four and five, and by 40% if between three and four. Moreover, in case the outer diameter of the isolator unit is less than 300 mm, a stress limit of 12 MPa shall be used for Category C. The isolator property variation shall not be greater than 20% of the nominal properties, and creep shall not exceed 5% of the total thickness.

•

Table 5.1 Compression Stress Limits of Rubber Isolator (GB50011, 2010) Building Category

Aa

Bb

Cc

Compression stress limit (MPa)

10

12

15

a

Major buildings that are not damaged during an earthquake event. Functional buildings that could be slightly damaged during an earthquake event. c Buildings that are not part of A and B. b

5.1 Available Codes and Design Tools

In addition to what is stated earlier, the isolator units shall meet the following additional requirements: •

•

Under rare earthquake occurrence, the isolation plane shall be stable and should not have permanent deformations. Moreover, the isolation units should not have any tension load. This is particularly difficult to be achieved for high-rise buildings; see Chapter 4 for same example applications. The stiffness (kd ) and damping (ζ d ) of the isolation plane can be computed as follows: X kd 5 kdi P kdi ζ di ζd 5 kd

•

•

(5.1) (5.2)

where kdi and ζ di are, respectively, the stiffness and equivalent viscous damping of the ith isolator unit that should be determined by testing. Horizontal stiffness shall be based on the shear force considered for rarely occurring earthquakes (2%
3% probability of exceeding the force within 50 years, as defined in GB-50011 (GB50011, 2010)). Torsional stiffness shall be taken into consideration in case torsional effects are present. The displacement of the ith isolator unit (udi ), under rare earthquake occurrence, should be limited as follows: udi 5 Ati udc # u i

(5.3)

where udc is the center of the isolation story displacement without torsion included, Ati is the ith isolator unit torsion factor (in the case of no torsion, a minimum torsion factor of 1.15 shall be used), u i is the displacement limit computed by the minimum between 0.55 times the effective diameter and 3 times the total thickness.

5.1.1.4.2 Buildings with energy dissipation devices GB 50011 (GB50011, 2010) allows two different types of devices: displacementdependent (metallic, friction, etc.) and velocity-dependent devices (viscous, viscoelastic, etc.). The following provisions are provided in the code: • • • •

•

Devices shall be provided at the story where the drifts are significant. Linear and nonlinear analyses can be utilized depending on the behavior of the energy dissipated system. A limiting value of 1=80 is required for the elasto-plastic interstory drift ratio. The device supplemental damping ratio can be estimated from the ratio between dissipated and potential energy as was shown in Eq. (3.60). The dissipated energy, for each device, can be computed similarly to what was shown in Chapter 4 and the potential energy as was shown in Eq. (3.61). The maximum supplemental damping is 20%.

245

246

CHAPTER 5 Design procedures for tall buildings

5.1.1.5 New Zealand code The New Zealand Concrete Structure Standard (NZS, 2006) provides simple guidelines for the seismic design of concrete structures equipped with added damping devices when utilized in base-isolated systems, which can also be applied to other structural systems. The standard states that the major advantages from inserting added damping devices as part of a base isolation approach can be gained in stiff structures rigidly connected to the ground (e.g., low-rise buildings or nuclear power plants). These devices are permitted if, at the ultimate limit states (ULS), the following requirements are satisfied: • • •

Device performance is supported by testing. Protection against the yielding of structural members is the same as the one considered for building without dissipation devices. Structural details allow the control of the deformation for demands greater than the design demands.

The design of the dissipation and main structural members should consider three earthquake intensity levels (and associated limit states): •

•

•

Moderate: the source of energy dissipation should be only from damping devices, that is, no damage to the structural members. This is considered a serviceability limit state (SLS). Design: the structural members could yield and typical standard design procedures (i.e., for structures without dampers) should apply, even if the structural system is subject to lower ductility demand as a result of the utilization of the dissipation devices. This is considered an ULS. Extreme: capacity design principles should be applied in order to prevent brittle failures and collapses. This is considered an extreme limit state.

The NZ standard (NZS, 2006) additionally suggests undertaking nonlinear time-history analyses (NLTHAs) for design purposes, since the knowledge and the experience in utilizing these systems are not well developed or established.

5.1.1.6 US codes In North America, the first attempt to apply a code to the design of structure with dissipation devices was undertaken in 1993 (Whittaker et al., 1991) by the Energy Dissipation Working Group (EDWG) of the Base-Isolation Subcommittee of the Structural Engineers Association of Northern California (SEAONC). Subsequently, in FEMA 222A (FEMA, 1994), NEHRP provided an appendix (not considered a code (Hanson and Miyamoto, 2002)) by introducing the utilization of these new techniques. The idea was to utilize an equivalent viscous damping approach with NLTHAs, when nonlinear velocity proportional devices were utilized. The first technical guideline available was Chapter 9 of FEMA 273 (FEMA, 1997a). This guideline was immediately superseded by the appendix to Chapter 13 of FEMA 302 (FEMA, 1997b). In this appendix, only general guidelines were provided.

5.1 Available Codes and Design Tools

A deeper insight in the utilization and design of energy dissipation devices were then provided in FEMA 356 (FEMA, 2000) and FEMA 450 (FEMA, 2003) (formerly FEMA 368 (FEMA, 2001)). These standards were a joint effort between ASCE and FEMA. The first document (FEMA, 2000) refers to the assessment of existing structures and the second (FEMA, 2003) examines the design of new structures. FEMA 450 (FEMA, 2003) was developed by Technical Subcommittee 12 (TS 12) of the Provision Update Committee (PUC). The proposed design philosophy was applicable to different dissipation systems (displacement- and velocity-dependent types). These standards specified the requirements on the type of analyses, design procedures, and testing protocols to utilize in the design process. The requirements of FEMA 450 were then included in ASCE 7-02 (ASCE, 2002) and subsequently in ASCE 7-05 (ASCE, 2005) and ASCE 7-10 (2010). A major revision has been just realized with ASCE 7-16 (ASCE, 2017a). This revision was based on the work conducted by TS 12, ASCE 41 committee members, and the NEHRP committee. The resulting work was based on the ASCE 41-13 (ASCE, 2013) and NEHPR (2015) recommendations and divided into two independent chapters: base isolated structures and structures with damping systems. In regard to existing buildings, TS 12 revised FEMA 356 (FEMA, 2000) in 2006 with ASCE 41-06 (ASCE, 2006) “Seismic Rehabilitation of Existing Buildings,” where guidelines for the seismic rehabilitation of structures with seismic isolation and energy dissipation devices are provided in Chapter 9. A major revision to this guideline was conducted in the 2013 edition of ASCE 41 (ASCE, 2013), where Chapter 13 goes into great detail for both seismic isolation and dissipative devices. The criteria selection between the two systems is based on the POs. Indeed, energy dissipation systems are usually preferred for several reasons: they are more suitable for a wider range of building heights; and they can also control other sources of excitation (like wind or mechanical loads). ASCE 41-13 (ASCE, 2013) also provides the possibility to utilize other control systems (e.g., active and hybrid control, and TMDs) as long as it is reviewed by an independent engineering review panel. However, no requirements for these devices are given. A new version of ASCE 41-17 (ASCE, 2017b) was just realized where, similar to ASCE 7-16 (ASCE, 2017a), base isolation and dissipation devices in two distinct chapters. Based on this historical review, in the proposed step-by-step procedures, the requirements of ASCE 7-16 (ASCE, 2017a), ASCE 41-17 (ASCE, 2017b), and NEHRP (2015) will be explained in detail with the relative design criteria that are the basis for the design of buildings equipped with dynamic modification technologies. These requirements will be applicable only to distributed dampers and base isolation systems, while for the other device categories the available literature will be reviewed for defining a design procedure. In addition, since the scope of this publication is on tall buildings, the requirements of ASCE (2017a) and NEHRP (2015) are integrated with the PEER/TBI (2017) and PEER/ATC (2010) guidelines that specify requirements for the performance-based seismic design (PBSD) of tall buildings.

247

248

CHAPTER 5 Design procedures for tall buildings

5.1.2 PRACTICAL DESIGN ASPECTS The design of tall buildings with additional dissipation devices requires the interaction of a number of professionals (wind tunnel laboratory, geotechnical engineer, damper manufacturer, owner, and architect), under the supervision of the structural designer. The professionals involved in the design process have different responsibilities from the characterization of the loads (wind, earthquake, vibrations), by means of numerical analyses or experimental activities, to the finalized design of the damping devices and the experimental characterization of all the mechanical components of the devices themselves and their behavior. During the structural analysis phase, specific assumptions are made by the structural engineer regarding the applied loads, the structural system, the boundary conditions, and the features of the damping devices; all of these modeling aspects, affecting the predictions of numerical analyses, are predefined and agreed upon together with the other professionals. Nevertheless, damping device characterization commonly requires an interactive process, involving different rounds of analyses with different sets of parameters. For this reason, it is important that wind tunnel tests, or the other analyses providing an estimate of the seismic action or vibration, shall be carried out in early stages in a way that they can provide reliable bases to modify the intensity of the external actions as a function of the considered return period and of the expected global damping. In fact, it is not uncommon for wind tunnel tests or seismic characterization studies to be carried out and completed well before the mechanical design of the damper is actually started, as an extensive interaction between the different professionals (i.e., climate and wind engineers, structural engineers, and manufacturers) is not always possible. In this scenario, the structural designer needs to make sure that, in the wind testing phase, the collected data provide tools for a robust evaluation of the relationship between the total damping ratio of the structure with additional devices (ζ T ) and the damping constant of the device (Cd ). Therefore, further interactions between the structural designer and the manufacturer can lead, in the following phase, to the finalized mechanical design of the devices.

5.1.2.1 Damper system from concept to production process Damping system installation in buildings is a straightforward process. Involvement of each stakeholder is necessary at the earliest design stage possible to understand the system as a whole. Strong interaction among the stakeholders plays a vital role in the success of the project. As soon as the building owner assigns an architect to design a building, architects must first consider whether this building will be built in a seismic and/or wind-prone area. This information will determine the shape and the dimension of building. They must also have general view of damping systems available in the market, so they can incorporate the appropriate system (if necessary) into their initial building design. Architects can seek advice from structural designers about what kind of damping system is suitable for their design. Both architects and

5.1 Available Codes and Design Tools

structural designers also need to give an explanation to the building owner about the importance of damping systems in a building to prevent unnecessary loss or damage. Building owners and architects must lead this crucial design stage. Certain aspects must be approved at the beginning of the design phase, such as additional cost, damper size, and location. The building owner can also assign other structural design firms to provide peer-review activities. Structural designers must gather as much information as possible from different damper manufacturers and check their track record and referral projects. Damper manufacturers need to provide their product specifications and required design parameters to the structural designer. Design assistance support by manufacturers will also help to verify good design practice for the structural designers. A detailed product specification is prepared based on good practice and international standards. However, the accuracy of input data is also very important. Correct measurement from wind laboratories, based on the proper model scale, will determine the quality of data. Site measurement will be required in order to understand the landscape and topography. This data will be a governing tool to make the decision on what type of damping system to use. Once the damping system is approved, it must be clearly specified through design drawings and tender documents, including a list of approved damper manufacturers. After tender approval by the owner, the damper is ready to be manufactured and shop drawings can be produced. These shop drawings will need to be checked thoroughly by the structural designer and building owner. Then, the damper is fabricated by the manufacturer. External supervision is necessary during this process to avoid any delay in schedule or defect in the product. Prototype and/or fabrication testing is done to check the accuracy and reliability before final delivery to site. Standard testing by an independent laboratory is also required to ensure the quality of the product (see Chapter 7 for more details about testing requirements for damping devices). During the construction stage, the contractor and structural designer must ensure that the damper installation follows the design assumptions. Installation must follow the standard procedure or method statement. On-site supervision is important in this case to minimize any mistakes in the installation and ensure optimal results. Final in situ measurements are recorded after installation to check the effectiveness of the damping system. See Chapter 7 for more details about inspection requirements.

5.1.2.2 Process of wind design and occupant comfort process Particularly for tall and slender buildings, occupant comfort (in the form of lateral accelerations) during wind events often controls the design of the lateral force
resisting system. The essential building characteristics under the control of the design team that affect a building’s response to wind include stiffness, mass, shape, and damping. The susceptibility of a building to experience large lateral accelerations may or may not be apparent to the structural engineer at the onset of the schematic

249

250

CHAPTER 5 Design procedures for tall buildings

design. Buildings with high height
width aspect ratios can be particularly challenging. For very slender buildings, it is often useful to engage the services of a wind tunnel laboratory during the schematic design phase. Performing an initial wind tunnel test can provide important information that can impact the design of the key elements during the schematic phase (e.g., the extent, location and size of the lateral force
resisting system components, the overall height of the building, or the shape of the tower). More often, with less challenging buildings, wind tunnel testing takes place during the design development phase, after the overall geometry of the building is set. However, stiffness, mass, and damping can still be adjusted to affect the building’s response. When performing either a high-frequency force balance study or a pressure integration study to determine the building response, the structural engineer will provide the wind tunnel laboratory with the building periods, in addition to the following properties for each floor of the building: • • • • •

Structural mode shapes Mass Location of the center of mass Mass moment of inertia Assumed level of intrinsic damping

The wind tunnel testing will yield results for equivalent static forces on a floor-by-floor basis, in addition to overall base shears and moments. Moreover, it is quite useful to produce spectra of along-wind and across-wind accelerations based on building period. The structural engineer will evaluate the results and iterate the process with the wind tunnel lab as necessary. During this process, a determination will be made as to whether a supplementary damping system needs to be incorporated into the design. Often, but not always, isolated damping systems are employed to mitigate wind effects. In such cases, the damper designer, and sometimes the wind tunnel lab, will be responsible for designing a system to provide the appropriate level of damping. From this point forward, the structural engineer merely has to keep the damper design apprised of any changes to the building dynamic properties as the design process moves toward completion.

5.1.2.3 Process of seismic design and strength requirements Whereas damping is generally considered in regard to serviceability issues, when designing for wind, damping affects strength requirements in seismic design. Obviously, based on the specific seismic demands, a building structure can be benefited greatly from higher levels of total damping. When the following prescriptive code approaches for the structural design, intrinsic levels of damping are built in to the code requirements (e.g., 2.5% ((ASCE 7-16 (ASCE, 2017a))), 3% (NEHRP, 2015)), as discussed in Section 3.2.

5.1 Available Codes and Design Tools

In cases where a PBD approach (PEER/TBI, 2017) is implemented using a supplementary damping system, the structural engineer must work together with the geotechnical engineer and damper manufacturer to develop the final design. A site-specific response spectrum may be developed, based on a range of total damping levels by the geotechnical engineer. The structural engineer can evaluate the lateral force
resisting system via modal response analysis for varied levels of seismic demand to understand the relative value of total damping. The geotechnical engineer will generally develop site-specific time-history ground motions to be incorporated in the structural engineer’s time-history analysis. A damper manufacturer can provide characteristics of specific damping elements to be incorporated into the analysis. While coordinating with the damper manufacturer, the structural engineer must iterate through his analysis to hone in on the specific characteristics required for the damping elements.

5.1.2.4 Properties of damping devices In order to correctly identify the basic parameters of damping devices, it is important to gather the best possible knowledge of the structural system and of all the factors affecting the efficiency of the dampers, such as construction tolerances, friction, additional stiffness, and temperature effects (see Section 5.1.3.3.3). The parameters used to define the constitutive law of dampers need to be specified and experimentally validated for different operational conditions, temperature ranges, and frequencies of excitation. In the case of dampers under wind loads, due to the nature of the load, all of the specifications for the materials used in manufacturing must be calibrated to take into account fatigue and deterioration issues. The service life of the device is specified and guaranteed by continuous monitoring and/or periodic inspection and maintenance; the maintenance plan is generally drafted by the manufacturer with the supervision of the structural designer (see Chapter 7 for further details).

5.1.3 STRUCTURAL ANALYSES Dynamic modification systems are holistic devices or systems, which serve to improve the performance of structures as a whole under dynamic excitation. By virtue of their height, tall buildings experience dynamic excitation caused by structure
wind interaction. Similarly, in geographic regions with relatively high seismicity, tall buildings are probably at risk of experiencing dynamic base excitation caused by intense ground shaking created by seismic events (earthquakes). Due to both of the aforementioned excitations, the response of the structure is primarily a function of mechanical properties unique to each building, such as natural frequencies, stiffness, mass, and damping. Input excitations themselves are functions of geometric profile/orientation, surface roughness, and geographic location (wind profile and seismicity), and act primarily in a lateral direction (parallel to the building’s floors).

251

252

CHAPTER 5 Design procedures for tall buildings

Consequently, an analytical framework, which approximates both the threedimensional (3D) dynamic properties of the structure and its site-specific loading history, must be employed when evaluating the implementation of dynamic modification systems. Commonly, this framework requires collaboration between different engineering consultants and is, by nature, iterative.

5.1.3.1 Computational structural analysis modeling Tall building systems inherently contain a large number of individual components acting together to form a lateral load
resisting system, which acts in opposition to dynamic excitation and loads created by wind and earthquakes. Taller buildings often have multiple lateral load
resisting typologies combined in a way to maximize lateral stiffness by engaging the individual stiffness of many structural members. Finite element analysis (FEA) techniques are well suited to determine the global, 3D, dynamic properties created by assemblies of large numbers of structural elements. Many modern commercial FEA software packages, specifically tailored to building analysis, exist and implement a variety of common analysis techniques/methodologies. Inherently, structural analysis models are mathematical approximations of the expected behavior of a physical system over time. Since buildings experience a complex loading history over their design life, the engineer must ensure that boundary condition sensitivity changes in system stiffness due to loading magnitude/displacement amplitude, material nonlinearity, and material property variability are accounted for. Therefore, the idea to bound the structural system’s response through multiple model runs should be employed. It is crucial that the engineer recognizes the approximate and iterative nature of structural analysis/modeling and, therefore, avoids attempting to capture complex behavior in one pass.

5.1.3.2 Analysis type The structural systems employed in tall buildings require the engineer to create multiple models when evaluating the performance of the structure. Different levels of analytical rigors are appropriate at different stages in the design process and models tend to evolve as design progresses and the structural system is refined. Common analytical methods (linear and nonlinear) are described in the following sections, in order of increasing rigor/complexity.

5.1.3.2.1 Linear analysis methods Linear analysis covers a broad range of methods, which use linear elastic material behavior to determine structural dynamic properties and responses. Linear analysis can be conducted using force-based approaches, such as linear static analysis and linear dynamic analysis (RSA); or it can be conducted using a linear timehistory approach (RHA). According to PEER/TBI (2017), linear analysis (RSA or RHA) is appropriate for service-level earthquake (SLE) evaluation and for design

5.1 Available Codes and Design Tools

earthquake (DE) when required by ASCE 7-16 (2017a). Some considerations for each linear approach are as follows: •

Linear static analysis is a design approach where equivalent static story forces, due to wind or earthquakes, are applied to the structure. The computation of story forces is prescriptive, and formulations for calculating these forces are provided within the applicable building code (Section 5.1.1). Linear static analysis is typically restricted to use in regular structures, where dynamic behavior is dominated by the fundamental mode of vibration, without significant higher modes and torsion effects and in regions of low seismicity. Since tall buildings often exhibit significant higher mode effects and the effects of torsion are important, linear dynamic analysis, instead of linear static analysis, is typically conducted for the seismic design of tall buildings, even in regions of low seismicity. • Linear dynamic analysis methods are based on procedures which employ the concept of modal superposition and are often associated with seismic design. Many building codes employ prescriptive provisions for modal RSA, where design acceleration response spectra are used to computate the peak linear response for each mode of vibration (Section 5.1.1). Peak modal responses are then statically combined (with a complete quadratic combination (CQC) method (PEER/TBI, 2017)) to generate the overall response of the structure. For wind-governed tall buildings, equivalent static wind forces provided by the wind consultant are often used in conjunction with linear dynamic/modal RSA for earthquakes. The structural engineer is not directly performing linear dynamic analysis for wind effects within their computation model, because the wind consultant creates the response spectra upon which the static loading provided is determined. A force spectrum using wind tunnel test data is used in conjunction with dynamic properties provided by the structural engineer to create wind response spectra in frequency domain, which are then used to create equivalent static loading and the associated load combination factors for x-sway, y-sway, and torsion. • Linear time-history analysis methods determine the structural response using inputs which vary with time. Acceleration versus time input signals can be applied to above-grade stories for determining response under wind loading, and base excitation accelerograms (earthquake records) can be used for determining response under seismic ground motion. Linear time-history analysis is typically conducted using modal analysis methods; whereas NLTHA can be conducted using fast nonlinear analysis (FNA, also called a modal method) or using direct-integration (DI) time-history analysis (as discussed in Section 5.1.3.2.3). Linear methods for dynamic modification systems. When force-based approaches, such as linear static or linear dynamic analysis, are used, it is helpful to think of results as a snapshot in time corresponding to a certain moment in the building’s loading history. Commonly, linear analysis is used to determine the

253

254

CHAPTER 5 Design procedures for tall buildings

baseline building properties provided to the wind consultant for use in motion assessment and to determine the equivalent static forces used for the calculation of interstory drift and other serviceability criteria. When motion-activated devices, such as TMD or tuned sloshing dampers, are employed, linear dynamic analysis is often utilized throughout the design process, with material properties varied when assessing the supplemental damping system’s performance. This approach is reasonable because motion-activated devices are most commonly employed to control accelerations resulting from dynamic wind interaction with the structure and are typically not used to improve seismic response. In regions of low seismicity, tall structures are designed to remain essentially elastic under service-level wind loading when compared to the much larger inelastic behavior observed under intense ground shaking due to earthquakes (in areas of high seismicity) and, therefore, a bounded linear dynamic analysis with appropriate assumptions/empirical parameters could be employed. In this scenario, modal RSA or linear time-history analysis for seismic design would generally follow code prescriptive methodology and would not account for any supplemental damping effects created under the ground motions. Although supplemental damping effect created by the device would not be used to improve response, its additional mass must be considered when determining seismic forces and building frequencies. Similarly, some displacement- or velocity-activated supplemental damping systems may be modeled using linear methods, provided that their behavior can be adequately captured using linear links (springs and dashpots). Whether linear or nonlinear analysis is used, damper link formations should be based on full-scale testing conducted by the damper manufacturer and must be carefully integrated into structural analysis models to ensure the link formulations are validated for the range of frequencies, displacements, and temperatures being investigated. If supplemental damping is ignored for earthquake response (i.e., wind serviceability only), the stiffness of the links must still be considered when determining seismic forces and building frequencies. It should be noted that when displacement- or velocity-activated distributed damping systems are approximated using linear analysis, supplemental damping produced is typically smeared over the entire model and equivalent modal damping is reported (for Eigen analysis to be employed, see the different step-bystep procedures as described in the later sections). When linear analysis methods are used, it is critical that appropriate effective element stiffness is considered for particular bounds, since, for example, concrete cracking, structural steel connection fixity, and foundation
soil interaction are not varied during a given model solution. Empirical stiffness modifiers are typically used to account for material nonlinear behavior for each analysis. The codes (e.g., ASCE, 2017a; PEER/ATC, 2010; AISC, 2016a; PEER/TBI, 2017; CSA, 2014) usually propose property modifiers for both material strength and effective stiffness, depending on the criteria the building needs to be designed for. Material strength shall be based on project-specific data. When these are not available, it is permissible to use expected properties instead of nominal or

5.1 Available Codes and Design Tools

Table 5.2 Expected Material Strengths (PEER/TBI, 2017) Material

Expected Yield Strength

Expected Ultimate Strength

482 MPa 565 MPa 476 MPa 586 MPa 1.1
1.5 fya

731 MPa 786 MPa 655 MPa 772 MPa 1.1
1.2 fua

Reinforcing Steel A615 Grade 60 A615 Grade 75 A706 Grade 60 A706 Grade 80 Structural Steel Concrete

1.3 fc

a

For structural steel the expected strength factors are different depending on the type of steel section utilized (i.e., hot rolled structural shapes and bars, hollow structural sections, steep pipe, and plates).

specified properties (see Table 5.2 for an example valid for US material (PEER/ TBI, 2017)). However, for evaluation and retrofit design of existing buildings, ASCE 41-13 (ASCE, 2013) specifies lower bound properties ( nominal properties) in force-based checks of elements that are prone to experiencing nonductile behavior, whereas it requires the use of “expected strength” when the element is expected to experience ductile behavior. Material stiffness shall take into account the effects of axial, flexural, and shear cracking. For steel members the elastic stiffness is determined by the full cross-sectional properties and the elastic modulus of steel (200,000 MPa) (PEER/ TBI, 2017). However, concrete member stiffness can be estimated with effective values as given in Table 5.3 (ASCE, 2017a), in Table 5.4 (PEER/TBI, 2017), and Table 5.5 (CSA, 2014), in lieu of detailed analysis. As it can be seen in the tables multipliers are less than one in most of the cases since they represent the effective linear branch of the inelastic model (PEER/TBI, 2017). In the tables, EC is the concrete elastic modulus; Ig is the cross-sectional gross moment of inertia; Aw is the cross-sectional web nominal area; and Ag is the cross-sectional gross area.

5.1.3.2.2 Nonlinear analysis methods Nonlinear analysis is a broad category of methods which use nonlinear material behavior (as well as geometrical) to determine the structural response. Depending on the specifics of a particular project, the level of detail of a nonlinear analysis can vary significantly. Where nonlinear analysis is used, it is common practice to begin analytical work with simpler linear models to establish baseline dynamic properties and global structural response. Understanding the basic system parameters and behavior will allow for informed decision making for what level of nonlinear analysis is warranted. As mentioned, modeling practices should be viewed as an evolution in rigor and detail. Multiple models with increasing levels of nonlinearity are typically employed until sufficient convergence is achieved.

255

256

CHAPTER 5 Design procedures for tall buildings

Table 5.3 Effective Stiffnesses of RC Structural Elements (for DE and MCER Demands) According to ASCE 41-17 (ASCE, 2017b) Component

Flexural Rigidity

Shear Rigidity

Axial Rigidity

Beams (nonprestressed)a Beams (prestressed)a Columnsb with compression 0 under gravity load $ 0:5Ag fc b Columns with compression 0 under gravity load # 0:1Ag fc or with tension

0:3Ec Ig Ec Ig 0:7Ec Ig

0:4Ec AW 0:4Ec AW 0:4Ec Aw



Ec Ag

0:3Ec Ig

0:4Ec Aw

Walls (cracked)

0:5Ec Ig

0:4Ec Aw

0:4Ec Ag

Ec Ag (compression) Es As (tension) Ec Ag (compression) Es As (tension) Ec Ag

0:4Ec Ag

Ec Ag

Flat slabs (nonprestressed) Flat slabs (prestressed)



   4c1 =l1 Ec Ig $ 1=3 Ec Ig

0:5Ec Ig

Note: c1 is size of the equivalent column in the direction of the span. l1 is the length of the slab in the direction of the seismic force. a For T-beams, Ig can be taken as twice the value of Ig of the web alone. For coupling beams in coupled shear walls, the effective stiffness values given for nonprestressed beams should be considered, unless alternative stiffness is determined by more detailed analysis (ASCE 41-17 (ASCE, 2017b)). b For columns with axial compression falling between the limits provided, flexural rigidity should be determined by linear interpolation. If interpolation is not performed, the more conservative effective stiffness should be used.

Note that this section is limited to discussion of the use of nonlinear elements for modeling components of the dynamic modification system itself. These nonlinear elements can be used in conjunction with an approximation of the main building structure which uses elastic elements. Bounding material properties of both the supplemental damping system and the main building structure to account for nonlinear behavior is an effective technique to integrate a dynamic modification system in structural engineering design. However, more advanced analysis can be undertaken where the main building structure is modeled using nonlinear elements. Nonlinear modeling of the total system, particularly for concrete structures, is very complex and is outside the scope of this book. Many of the seismic references cited in this publication offer limited guidance on how to appropriately undertake these analyses. Moreover, it is important to note that whenever undertaking comprehensive nonlinear analysis of the total building system, peer review should be undertaken. When displacement- or velocity-activated supplemental damping systems, such as metallic dampers, friction dampers, viscous damping devices, and viscoelastic damping devices, or base isolation are employed, nonlinear analysis is often utilized because it can capture the behavior of dampers modeled using link elements. Moreover, two common nonlinear time-history methodologies can be utilized: FNA and DI analysis. FNA is more commonly employed when modeling

Table 5.4 Effective Stiffnesses of RC Structural Elements According to PEER/TBI (2017) MCER-Level Nonlinear Models

Service-Level Linear Model Component a

Structural walls (in-plane) Structural walls (out-of-plane) Basement walls (in-plane) Basement walls (out-of-plane) Coupling beams with conventional or diagonal reinforcement Composite steel/reinforced concrete coupling beamsb Non-post-tension (PT) transfer diaphragms (in-plane only) PT transfer diaphragms (in-plane only) Beams Columns Mat (in-plane) Mat (out-of-plane) a

Axial

Flexural

Shear

Axial

Flexural

Shear

1:0Ec Ag
1:0Ec Ag
1:0Ec Ag

0:75Ec Ig 0:25Ec Ig 1:0Ec Ig 0:25Ec Ig  0:07 hl Ec Ig # 0:3Ec Ig

0:4Ec Ag
0:4Ec Ag
0:4Ec Ag

0:75Ec Ag
1:0Ec Ag
1:0Ec Ag

0:35Ec Ig 0:25Ec Ig 0:8Ec Ig 0:25Ec Ig  0:07 hl Ec Ig # 0:3Ec Ig

0:2Ec Ag
0:2Ec Ag
0:4Ec Ag

1:0ðEAÞtrans

0:07

1:0Es Asw

1:0ðEAÞtrans

0:07

0:5Ec Ag

0:5Ec Ig

0:4Ec Ag

0:25Ec Ag

0:25Ec Ig

0:1Ec Ag

0:8Ec Ag 1:0Ec Ag 1:0Ec Ag 0:8Ec Ag


0:8Ec Ig 0:5Ec Ig 0:7Ec Ig 0:8Ec Ig 0:8Ec Ig

0:4Ec Ag 0:4Ec Ag 0:4Ec Ag 0:8Ec Ag


0:5Ec Ag 1:0Ec Ag 1:0Ec Ag 0:5Ec Ag


0:5Ec Ig 0:3Ec Ig 0:7Ec Ig 0:5Ec Ig 0:5Ec Ig

0:2Ec Ag 0:4Ec Ag 0:4Ec Ag 0:5Ec Ag


When modeled as line elements and not as fiber elements. ðEIÞtrans is the cracked trasformed sections.

b

l h

ðEIÞtrans

l h ðEIÞtrans

1:0Es Asw

258

CHAPTER 5 Design procedures for tall buildings

Table 5.5 Effective Stiffnesses of RC Structural Elements According to Canadian Standard Association (CSA) A23.3 (CSA, 2004) Ranges for Linear Elastic Analysis Load Type Wind

Component Nondiagonally reinforced coupling beams Diagonally reinforced coupling beams Shear walls Shear walls in net tension Slabs with mild reinforcement Posttensioned slabs Beams (excluding coupling beams) Columns Columns in net tension

Seismic

Beams (excluding coupling beams) Columns (Nondiagonally reinforced) coupling beams (Diagonally reinforced) coupling beams Slab frame element Wall

Serviceability (SLS)

Strength (ULS)

0:50Ig Ave 5 0:40Ag 0:45Ig Ave 5 0:45Ag 0:95Ig Ave 5 0:95Ag

0:40Ig Ave 5 0:25Ag 0:35Ig Ave 5 0:40Ag 0:75Ig Ave 5 0:75Ag

Refined calculation required 0:35Ig 0:60Ig 0:50Ig Ave 5 0:75Ag 1:0Ig Ave 5 1:0Ag

0:20Ig 0:45Ig 0:40Ig Ave 5 0:50Ag 0:70Ig Ave 5 0:70Ag

Refined calculation required 0:40Ig a c Ig 0:40Ig Ave 5 0:15Ag 0:25Ig Ave 5 0:45Ag 0:20Ig aw Ig Axe 5 aw Ag

0:40Ig ac Ig 0:40Ig Ave 5 0:15Ag 0:25Ig Ave 5 0:45Ag 0:20Ig aw Ig Axe 5 aw Ag

Note: From CSA A23.3 (CSA, 2014) Table N9.2.1.2 and Table 21.1.

distributed supplemental damping systems because it is found to be generally more accurate and efficient than DI methods. Additionally, typically only FNA allows for the generation of energy plots which can be used to evaluate the performance/damping generated by link elements. However, if the model has many DOFs and large number of link elements are utilized, DI may be more straightforward and efficient than FNA (this is true if limited number of response-history runs is expected). Moreover, FNA solver do not provide the nonlinear geometry—large displacement option and such effects cannot be studied unless DI methods are utilized.

5.1 Available Codes and Design Tools

When modeling distributed supplemental damping systems, additional modes of vibration need to be considered to capture the damping produced by (nonlinear) link elements. Ritz vectors require much fewer additional modes of vibration to capture link damping than eigenvectors and therefore it is commonly recommended to use them. Further, attempting to use eigenvectors with nonlinear link elements may result in capturing the incomplete or inaccurate behavior since nonlinear analysis does not smear damping globally. Therefore, it is strongly recommended that before undertaking complex structural analysis which integrates the effects of a supplemental damping system, the engineer should thoroughly review the various analytical methods available. When conducting FNA using Ritz vectors, as recommended, an additional mode of vibration must be provided for each nonlinear link used in the model. In general, prior to implementing a model with nonlinear links, the engineer will have run baseline models to determine dynamic properties ignoring the effects of the dynamic modification system. This baseline model should be used to determine the minimum number of modes required to achieve appropriate amount of mass participation (a requirement in many building codes for seismic design). Once the number of modes required to quantify global building behavior is established, additional modes should be included to capture each nonlinear link in the supplemental damping system. Convergence studies should be employed to ensure appropriate number of modes are provided and the hysteretic plots of each link should be reviewed to ensure the links are active in the model. Since FNA does not report lumped equivalent modal damping, log-decrement displacement versus time plots are commonly used to compute supplemental damping. As recommended previously, models should evolve in complexity as the behavior of the system is better understood/refined. For distributed supplement damping systems, linear links with eigenvectors can be used to provide a preliminary (unconservative) upper bound of the supplemental damping provided. When more advanced nonlinear links are used with Ritz vectors, the supplemental damping produced in a distributed system should be marginally lower than predicted using the linear links. Caution: Where the supplemental damping system is not a distributed system, but is rather concentrated in a small number of elements, linear analysis with eigenvectors will not provide realistic preliminary level results and the engineer should begin by using Ritz vectors. Damped outriggers or belt trusses would be an example of such a concentrated system, whereas distributed viscoelastic coupling dampers replacing a significant number of coupling beams is an example of a “distributed” system. Caution: Related with “damping leakage” phenomenon with FNA. This effect produces an overestimation of damping when high values of linear effective are utilized in certain software (Sarlis and Constantinou, 2010). Common computational methods for calculating supplemental damping. Classical linear modal analysis with eigenvectors is a classical modal

259

260

CHAPTER 5 Design procedures for tall buildings

analysis and assumes that the damping provided by the dynamic modification system is proportional to velocity. If the damping matrix can be decoupled using the mode shapes, the mth modal damping coefficient can be obtained by pre- and postmultiplying the damping matrix by the mth mode shape. The damping ratio in the mth mode is obtained as a function of the mth circular frequency and modal mass. Most commercial software packages will report damping values mode by mode numerically in their output files. As noted previously, this technique can overestimate or underestimate the added damping; however, it is accurate if the damping matrix can be decoupled using the mode shapes. It is therefore recommended to be used only for preliminary analysis wherever appropriate (as noted previously). Nonlinear modal analysis with Ritz vectors requires the use of free vibration study, which is produced by a ramp load followed by an instantaneous removal of the load (this is valid for SDOF since for MDOF the shape of load is also important). Structural response output data are used to calculate damping using the logdecrement technique as the average of the ratio of the positive peak and negative peak amplitude of each peak followed by the mth following cycle. It is common to discard the first several damping value estimations because they are typically more variable (since the response get more stable after a couple of cycles). Subsequently, the total damping is taken as the average of remaining damping values. Supplemental damping matrix is determined by subtracting the inherent damping matrix, which is commonly an input parameter in FEA software, from the total computed damping matrix. P-delta analysis. Modern commercial FEA software packages used for the lateral analysis of buildings have features/functions which include P-delta effects. The destabilizing gravity loads shall be considered for P-delta effects over the entire building. In case only the lateral resisting system is modeled, leaning columns shall be considered to provide realistic gravity load distribution (PEER/ATC, 2010; PEER/TBI, 2017).

5.1.3.2.3 Preliminary analyses Although static and dynamic analyses of tall building structural systems can be nowadays be performed with the aid of commercial finite element packages, in early stages of design, it is still a common practice to adopt simplified models. For example, the adoption of continuum models, the so-called replacement beam (RB) models, can help designers to more properly deal with the analysis and preliminary design of building structures. The dominant behavior of tall building structural systems can be captured with clamped (cantilever) RB models (Euler Bernoulli beam (EBB), shear beam, Timoshenko beam, coupled two beams, and sandwich beam (SWB)) by considering the stiffness properties (Fig. 5.7 (Faridani, 2015)). The damping phenomenon can also be studied by means of distributedparameter models (Lavan, 2012; Faridani and Capsoni, 2016a,b). In order to take into account the additional damping, distributed-parameter damping formulations

5.1 Available Codes and Design Tools

FIGURE 5.7 Damper synthesis of the RB idealization process from the 3D model to a 1D model.

have been developed for the EBB frame to approximately analyze shear wall systems including supplementary viscous damping (Lavan, 2012).

5.1.3.3 Boundary conditions and common assumptions When undertaking global structural analyses, it is often necessary to perform sensitivity analysis to bound uncertainty/variability using several different models. The approach of bounding the probable behavior of a structure when approximating it in FEA should also be employed when integrating damper elements in analysis. As structural complexity increases, determining the governing criteria globally and locally becomes increasingly difficult, and it therefore becomes critical to be able to process large volumes of data efficiently. Adopting a framework of automated postprocessing of data allows the engineer to quickly determine the sensitivity of selected variables. The following subsections briefly discuss some potential parameters which commonly require sensitivity analysis/bounding.

5.1.3.3.1 Soil
structure interaction Soil
structure interaction (SSI) is commonly accounted for by a bounded analysis employing several different models. Where floor plan enlarges below grade a virtual outrigger is created by foundation walls engaged laterally by slab diaphragms, and this virtual outrigger action is commonly referred to as the “backstay effect” (PEER/ATC, 2010). Backstay effects are particularly sensitive to diaphragm rigidity and to the surrounding soil interacting with the foundation walls. It is common to run upper and lower bounds of slab diaphragm stiffness to conservatively envelope forces in the diaphragms (collectors), foundation walls, and the tower’s primary lateral load
resisting system as it passes through the below-grade structure.

261

262

CHAPTER 5 Design procedures for tall buildings

Modeling SSI and bounding backstay effects are outside the scope of this book but are thoroughly covered in PEER ATC-72-1 (PEER/ATC, 2010) (Appendix A), PEER Guidelines for Performance-Based Design of Tall Buildings (PEER/TBI, 2017), and LATBSDC Alternative Analysis and Design Procedure (LATBSDC, 2017).

5.1.3.3.2 Effective stiffness for RC components Effective stiffness parameters for RC components are commonly bounded to account for increased cracking at high-load levels (ULS) and for lower cracking at service load levels used for wind serviceability (drifts and building motions). As shown earlier, Tables 5.3
5.5 provide guidance on concrete cracking-related stiffness parameters for various elements and different demand levels (SLE, DE, and maximum considered earthquake (MCER)). Typically, a supplemental damping system’s efficiency is evaluated at both service and ultimate load levels; and for seismic design for a wide range of ground motion intensities when conducting PBSD. Generally cracking modification factors are employed in linear elastic models to obtain baseline structural dynamic properties (from FEA). Sensitivity analysis of high-demand RC elements can be conducted by extracting forces and rotations from a linear elastic model and employing nonlinear concrete models such as modified compression field theory (Vecchio and Collins, 1986) or other constitutive models to determine cracking at a particular load level and updating effective stiffness parameters in linear FEA models iteratively, thereby creating a quasi-nonlinear model at that load level. This approach creates discrete models for a particular time in a structure’s loading time history. For advanced analysis, particularly in the high seismicity area, nonlinear timehistory FEA is often used when modeling the total building system at relatively high-load levels. Since these models are complex and tend to have relatively long analysis time, bounding effective stiffness in linear elastic or quasi-nonlinear models is an effective way to begin to understand the behavior of the system prior to running full nonlinear models. Moreover, for the design of wind-governed supplemental damping systems/in areas of low seismicity, it is often common practice to bound concrete nonlinear material properties in quasi-nonlinear models. When bounding the effects of concrete cracking, particular attention should be paid to high-demand elements and concentrated stiffness systems such as: • • • • •

Coupling beams Outrigger walls Belt walls Columns used to couple outriggers and belts to the rest of the lateral load
resisting system Diaphragms/collectors engaging secondary lateral load
resisting elements/ systems (such as foundation walls)

5.1 Available Codes and Design Tools

Other mechanical and material properties which may require the use of bounding/sensitivity analysis: • • •

•

Young’s modulus of concrete Joint rigidity and fixity in steel and concrete structures Foundation bearing strata stiffness; particularly on soft strata: • Raft foundation rotation • Pile
soil interaction Effect of secondary elements

Note: secondary elements such as slabs (out-of-plane flexural rigidity), gravity beams, and continuous drop panels may affect the lateral stiffness and frequencies of the tower. However, the stiffness of these elements is often ignored when determining the demand on the primary lateral load
resisting system. In this approach, the design of the secondary elements would consider gravity and lateral forces (determined from a model where their stiffness is not ignored).

5.1.3.3.3 Modification factors of dynamic modification properties Nominal properties of dynamic modification system shall take into account the variation due to different factors, such as temperature, aging, and environmental exposure (ASCE, 2017a; NEHRP, 2015). In general, to account for property variation factors in modeling, there are three methods: 1. Using advanced software packages which explicitly models property variation due to environmental conditions (e.g., temperature and velocity). More detail about this approach is outside the scope of this book. 2. Through testing data (see Chapter 7). 3. Use of property modification factors (λ) (see subsequent discussion). Property modifications factors (λ) should be defined for each category of dynamic modification devices. US codes provide different expressions to calculate them: •

NEHRP (2015) and ASCE 7-16 (ASCE, 2017a) recommend that a minimum, λmin , and a maximum, λmax , variation factor shall be employed in damper modeling, analysis, and design. These two factors are functions of minimum and maximum values of three parameters: aging/environmental factors (λae ), testing (λtest ), and specifications (λspec ), as follows:     λmax 5 1 1 0:75 λae;max 2 1 λtest;max 3 λspec;max $ λmax;lim     λmin 5 1 2 0:75 12λae;min λtest;min 3 λspec;min # λmin;lim

(5.4) (5.5)

Practical ranges are (NEHRP, 2015; ASCE, 2017a): 0:85 # λmin # 0:95 and 1:05 # λmax # 1:15. The limiting values for distributed dampers are: λmax;lim 5 1.2 and λmin;lim 5 0.85; while for base isolation the following values are recommended: λmax;lim = 1.8 and λmin;lim 5 0.8 or 0.6 for NEHRP (2015) and ASCE (2017a) respectively.

263

264

CHAPTER 5 Design procedures for tall buildings

Maximum and minimum design properties can be obtained by multiplying nominal design properties of dampers by the property variation factors λmin and λmax , as follows:

•

Maximum property 5 Nominal property 3 λmax

(5.6)

Minimum property 5 Nominal property 3 λmin

(5.7)

Then, separate analyses and design for the building with maximum and minimum properties should be performed under external excitation (e.g., wind and SLE/DE/MCER earthquake conditions). ASCE 41-17 (2017b): similar to NEHRP (2015) and ASCE (2017a) recommendations the minimum, λmin , and maximum, λmax , modification factors can be computed as follows:     λmax 5 1 1 SPAF λPM;max 2 1 λspec;max $ λmax;lim     λmin 5 1 2 SPAF 12λPM;min λspec;min # λmin;lim

(5.8) (5.9)

where λspec is the ower and upper specification tolerance, which is the permissible variation between the average production test values and the nominal values (in the range of 1 10% to 1 15%); λPM is the global property modification factor, which is the product of all the factors for environmental and testing effects (typical values are provided in ASCE (2017a) based on AASHTO recommendations (AASHTO, 2010)); SPAF is the system property adjustment factor, which takes into account that each property modification factor does not occur at the same time (taken as 0.67 for all performance levels). ASCE 41-17 (ASCE, 2017b) provides guidance on this factor based on AASHTO recommendations (AASHTO, 2010); λmax;lim is the maximum limit value equals to 1.15 or 1.3 for distributed dampers and base isolation, respectively; and λmin;lim is the minimum limit value equals to 0.85 for both distributed dampers and base isolation. Multiplying nominal design properties of dampers by property variation factors λmin and λmax leads to maximum (Eq. 5.6) and minimum (Eq. 5.7) design properties.

5.2 PASSIVE DAMPING SYSTEMS Including passive damping systems as part of a structural design scheme is based on a well-maintained balance between performance targets, construction costs, and complexities of their implementation in an actual building. Every tall building or dynamically sensitive structure has unique characteristics and constraints that must be carefully investigated by the design team in order to determine which

5.2 Passive Damping Systems

type of damping system can best achieve the target performance in terms of cost, construction schedule, constructability, etc. The up-front implementation assessment of a dynamic modification system should consider a wide range of factors that are critical in achieving the desired performance, such as: • • • • • • • • •

Source of the external dynamic excitation (e.g., vortex shedding, turbulent buffeting, seismic) Anticipated dynamic behavior of the structure Performance level required based on the occupancy or usage Damper
structure interaction Load magnitude at the damper
structure interface Available space Construction material, method, and schedule Lifting capabilities at the site during installation Maintenance and inspection requirements

At early design stages, careful consideration of the above factors are fundamental to the successful implementation of dynamic modification systems, and to assure that the desired performance of the building is achieved. As shown in the previous chapters, within the category of the passive damping systems, mass damping systems (i.e., TMDs and TLDs) have become increasingly popular solutions to enhance the serviceability performance of tall buildings and other wind-sensitive structures. However, the implementation of distributed damping systems (i.e., viscoelastic dampers, viscous dampers, and friction dampers) has been primarily aimed at providing earthquake protection for structures in high seismic regions. Basic design considerations and practical suggestions for implementing these systems are described with a step-by-step procedure (Fig. 5.8) for the different damping devices discussed in Chapter 4. Ten main steps are defined (Fig. 5.8) and detailed explanations of each step are given in the following sections.

5.2.1 STEP-BY-STEP PROCEDURE FOR DISTRIBUTED DAMPERS In Section 4.1.1, the basics principles of design for distributed dampers were introduced. In this section, the design procedure for a building with distributed dampers will be reviewed. In the literature, there are several methods available for both code/guidelines and noncode/guidelines prescriptive procedures, as summarized in Table 5.6. The majority of these procedures refer to design of structures under seismic loading without dealing with wind design (except the work by McNamara et al. (1999)). The main reason for this could be related to the dominance of the design procedures based on low-rise buildings that usually are dominated by seismic excitations. Indeed, while looking at tall buildings, wind loading can become the predominant concern as shown in Chapter 3. Another important consideration is related to the type of analyses allowed by these procedures. Standard building

265

266

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.8 Step-by-step procedure for implementing passive damping systems.

5.2 Passive Damping Systems

Table 5.6 Distributed Damper: Available Design Procedures Code/Guidelines Prescriptive

Noncode/Guidelines Prescriptive

FEMA-356 (FEMA, 2000d) FEMA-368 (FEMA, 2001) FEMA-450 (FEMA, 2003) ASCE 41-13 (ASCE, 2013) ASCE 41-17 (ASCE, 2017b) ASCE 7-10 (ASCE, 2010) ASCE 7-16 (ASCE, 2017a) TBI (PEER/TBI, 2017) FEMA P-751 (NEHRP, 2012) FEMA P-1050-1 (NEHRP, 2015) Ramirez et al. (2001) Liang et al. (2012) Chinese Code (GB50011, 2010) Japanese Code (BCJ, 2013) JSSI Manual (2003, 2005, 2007)

Gluck et al. (1996) Soong and Dargush (1997) Peckan et al. (1999a,b) Lopez-Garcia (2001) Yang et al. (2002a,b) Kim et al. (2003) Uetani et al. (2003) Lin et al. (2003) Christopoulos and Filiatrault (2006) Lin et al. (2008) Lomiento et al. (2010) Silvestri et al. (2010) Lago (2011); Sullivan and Lago (2012) Zhou et al. (2012) Pettinga et al. (2013) Guo and Christopoulos (2013) Diotallevi et al. (2014) Palermo et al. (2016)

codes, such as the Chinese (GB50011, 2010), Japanese (BCJ, 2013), and US codes (ASCE, 2017a), permit the use simplified analysis procedures only for regular and low-rise buildings and, in the case of tall buildings, more complicated approaches are required, including dynamic analyses (for seismic design) and wind-tunnel tests (for wind design) (e.g., ASCE 7-16 (ASCE, 2017a) and PEER/TBI (2017)). Given the scope of this book, only the most reliable and generally recognized design approaches are examined. Code-prescriptive procedures are the most suitable procedures for design, since they are based on current national standards. For these reasons, in the following sections, the step-by-step procedure provided is mostly based on US design codes (ASCE 7-16 (ASCE, 2017a), ASCE 41-17 (ASCE, 2017b), NEHRP (2009, 2015), PEER/TBI (2017)). However, reference to other methods is given to underline the current trends in the research field. The procedure has been developed for both seismic and wind design. Moreover, this approach has been made as broad as possible, so that it can be applicable to different types of distributed dampers (e.g., viscous, viscoelastic, metallic, and friction) and building heights. In the following sections, every step, as shown in Fig. 5.8, is developed in detail for four different types of damping devices: viscous, viscoelastic, metallic (hysteretic), and friction.

267

268

CHAPTER 5 Design procedures for tall buildings

Table 5.7 Risk Category of Tall Buildings for Flood, Wind, Snow, Earthquake, and Ice Loads (ASCE, 2017a) Importance Factors Occupancy of Tall Buildings Buildings that represent a low risk to human life in the event of failure All buildings except those listed in risk categories I, III, and IV Buildings which could present a significant risk to human life under failure Buildings designated as essential facilities

Risk Category

Earthquake Loads (Ie )

Ice-Wind Loads (Iw )

I

1

1

II

1

1

III

1.25

1

IV

1.5

1

5.2.1.1 Step 1: Building and site categorization This step is subdivided into three subsections. The first subsection (Step 1.1) is devoted to the categorization of tall building sites in which the building is located. The second subsection (Step 1.2) relates only to buildings that need to be designed for seismic activity and it differentiates between the code prescriptive RSA and ground motion hazard analysis for the MCER. The third subsection (Step 1.3) is devoted to the determination of basic wind design loads.

5.2.1.1.1 Step 1.1: Risk category and occupancy importance factor Buildings are categorized to determine flood, wind, snow, ice, and earthquake loads, based on the risk associated with different performance levels (ASCE, 2017a). The building risk category and relative occupancy importance factor can be found using Table 5.7, depending on the priority of minimum design loads (e.g., seismic and wind) for the building. These values will be used to determine the seismic and wind loading that the building needs to sustain, as is shown in subsequent steps.

5.2.1.1.2 Step 1.2: Site spectral response acceleration, response spectrum, and time histories For the seismic design of buildings with dynamic modification systems, different seismic levels are considered (ASCE 7-16 (ASCE, 2017a), NEHRP (2015)): (1) DE and (2) MCER. The responses to the first set are suggested (NEHRP 2015) to be used for the design of structural systems, equipped with dynamic modification devices, and the responses to the second set are to be employed for the design of dynamic modification systems (ASCE, 2010; NEHRP, 2015). However, for tall buildings the TBI (PEER, 2017) recommends to use a PBD procedure according to the following two seismic levels: SLE (instead of DE) and MCER. Tall buildings as per PEER (2017) have the following characteristics: • •

Fundamental period greatly above 1 second. Higher mass participation for higher modes of vibration.

5.2 Passive Damping Systems

• •

Drift ratio from axial (from walls and columns) and shear (from frames or walls) deformation is comparable. Slender aspect ratio for the seismic force-resisting system.

The seismic demand on a building is based on the risk-targeted MCER, a site characteristic, of where the building will be constructed. According to ASCE 716 (ASCE, 2017a) the spectra for both DE and MCER levels can be determined according to the following procedures depending on the MCER spectral response acceleration at 1 second (S1 ): • •

If S1 , 0:6, consider an MCER based on mapped values and determine the corresponding spectra for the site DE and MCER levels. If S1 $ 0:6, perform an MCER ground motion hazard analysis in accordance with Section 21.2 of ASCE 7-16 (ASCE, 2017a) and then construct the relevant spectra for the site DE and MCER levels.

Buildings with risk categories I, II, III, or IV, located in a site where S1 $ 0:75, should not be constructed, where there is a known potential for an active fault to cause a rupture in the ground surface at the structure. The spectral acceleration for both the short period (SS ) and at 1 second (S1 ), given the site location, should be computed according to one of the following methods: 1. Directly determined using Figs. 22-1
22-6, presented in ASCE 7-16 (ASCE, 2017a). These values refer to site class B and for other soil site class values shall be modified through correction factors (ASCE, 2017a). 2. Using the online tool available through the USGS website. Note: this S1 categorization is not valid for tall buildings since for ground motion characterization only hazard analysis is allowed (i.e., mapped values cannot be considered) (PEER, 2017). However, in the following both mapped values and hazard analysis are reviewed for completeness. S1 , 0:6 Risk-targeted maximum considered earthquake (MCER)-based on mapped values (not applicable to high-rise buildings according to PEER (2017)). According to ASCE (2017a), the response spectrum is constructed based on the following five steps: 1. Site class: Based on the site-soil properties, the site can be classified into six categories A, B, C, D, E, or F. The most appropriate parameter that determines the site class is the average value of the shear-wave velocity, vs , of the upper 100 ft (30.48 m) of soil at the site, which is expressed as follows: Pn di vs 5 Pi51 n di

(5.10)

i51 vsi

Here, n is the number of layers of similar soil materials, where data are available; di is the depth of layer i; and vsi is the shear-wave velocity of the soil in n P layer i. It should be noted that di equals 100 ft (30.48 m). i51

269

270

CHAPTER 5 Design procedures for tall buildings

Table 5.8 Site Class Classification Based on Site Material and Average Shear Wave Velocity v s (ASCE, 2017a) Site Class A B C D E F

Range of v s (m/s)

Site Material Hard rock Rock Very dense soil and soft rock Stiff soil Any profile with more than 3 m of soft clay Liquefiable soils; quick and highly sensitive clays; collapsible weakly cemented soils; peats or highly organic clays; very high plasticity clays; very thick soft or medium-stiff clays

.1524 762
1524 366
762 183
366 ,183 _

Having determined v s using Eq. (5.10), the site classes can be easily selected from Table 5.8. According to ASCE 7-16 (ASCE, 2017a) there are two exceptions when using Eq. (5.10): a. If the average value of shear-wave velocity, v s , is not known for a site, two other parameters, the standard penetration test blow count, N , and the undrained shear strength, s u , of the upper 100 ft (30.48 m) of soil at the site, can be applied. b. If the soil properties are not known in sufficient detail to determine the site class, class D can be used, unless the expert consultant (geotechnical engineer) determines that site class E or F is existing at the site. 2. Design spectral response acceleration parameters: The design spectral response acceleration parameters (SDS and SD1 ), at short period, TS , and at a 1-second period, T1 , can be estimated as follows: 2 SDS 5 Fa SS 3

(5.11)

2 SD1 5 Fv S1 3

(5.12)

Here, Fa and FV are site coefficients, which are defined in Tables 5.9 and 5.10, respectively. To determine these coefficients, site class and mapped MCER spectral response acceleration parameters (SS and S1 ) are required (ASCE, 2017a). c. Seismic design category: The building can be assigned to a seismic design category (use Table 5.11) based upon their risk category and the severity of the DE at the site (i.e., design spectral response acceleration parameters, SDS and SD1 , determined previously). d. Design response spectrum: If site-specific ground motion procedures are not employed, a design response-spectrum curve can be developed (see Fig. 5.9) as:

5.2 Passive Damping Systems

Table 5.9 Values of Site Coefficient Fa (ASCE, 2017a) Range of SS a

Site Class

SS # 0.25

SS 5 0.5

SS 5 0.75

SS 5 1.0

SS 5 1.25

SS $ 1.5

A B C D E

0.8 0.9 1.3 1.6 2.4

0.8 0.9 1.3 1.4 1.7

0.8 0.9 1.2 1.2 1.3

F

Site analysis required

Site analysis required

Site analysis required

0.8 0.9 1.2 1.1 Site analysis required Site analysis required

0.8 0.9 1.2 1.0 Site analysis required Site analysis required

0.8 0.9 1.2 1.0 Site analysis required Site analysis required

a

Use straight-line interpolation for intermediate values of Ss .

Table 5.10 Values of Site Coefficient Fv (ASCE, 2017a) Site Class A B C D E

F

a

Range of S1 a S1 # 0.1

S1 5 0.2

S1 5 0.3

S1 5 0.4

S1 5 0.5

S1 $ 0.6

0.8 0.8 1.5 2.4 4.2

0.8 0.8 1.5 2.2b Site analysis required Site analysis required

0.8 0.8 1.5 2.0b Site analysis required Site analysis required

0.8 0.8 1.5 1.9b Site analysis required Site analysis required

0.8 0.8 1.5 1.8b Site analysis required Site analysis required

0.8 0.8 1.4 1.7b Site analysis required Site analysis required

Site analysis required

Use straight-line interpolation for intermediate values of S1 . Also see requirements for site analysis.

b

SA 5

0 1 8 > > T > > SDS @0:4 1 0:6 A for T , T0 > > T0 > > > > > > S for T $ T0 ; T # TS > DS < SD1 > > > > > > T > > > SD1 TL > > > > : T2

for T . TS ; T # TL

(5.13)

for T . TL

where T is the fundamental period of the structure; T0 equals to 0:2ðSD1 =SDS Þ; TS equals to SD1 =SDS ; and TL is the long-period transition parameter obtained from published maps on ASCE 7-16 (ASCE, 2017a), or site-specific response

271

272

CHAPTER 5 Design procedures for tall buildings

Table 5.11 Seismic Design Category Based on SDS and SD1 (ASCE, 2017a) Risk Category Range of SD

I or II or III

IV

SDS , 0.167 or SD1 , 0.067 0.167 # SDS , 0.33 or 0.067 # SD1 , 0.133 0.33 # SDS , 0.5 or 0.133 # SD1 , 0.2 SDS $ 0.5 or SD1 $ 0.2

A B C D

A C D D

FIGURE 5.9 Design response spectrum.

analysis, or any other methods approved by the authority having jurisdiction. The details of the procedure and rationale used in determining the TL maps in ASCE 7-16 (ASCE, 2017a) and in NEHRP (2009) can be found in Crouse et al. (2006). e. Risk-targeted maximum considered response spectrum: Given the design response spectrum (Fig. 5.9), the MCER response spectrum can be simply determined by multiplying the design response spectrum by 1.5 (ASCE, 2017a; NEHRP, 2015). S1 $ 0:6 Risk-targeted maximum considered earthquake (MCER) ground motion hazard analysis (ASCE, 2017a) (for any value of S1 for tall buildings according to PEER (2017)). An hazard analysis shall be conducted for tall buildings (according to PEER (2017)) and for any other structures in the case S1 , previously obtained, is greater than 0.6 second. This shall account for the geological and seismic characteristic of the site to be provided in a report (ASCE, 2017a). The site-specific MCER is determined from two different procedures: (1) probabilistic and (2) deterministic ground motions. Based on these, the site-specific

5.2 Passive Damping Systems

spectral response accelerations (SAM ) shall be defined, at any period, as the lesser of the values computed from the probabilistic and deterministic methods. In the following, the two different procedures are briefly reviewed: 1. Probabilistic risk-targeted MCER ground motions: the response spectrum for this method has a return period of 1% probability of collapse within a 50-year period. The spectral acceleration coordinate can be determined from either of the two following methods: a. The acceleration at each spectral response period can be computed from the spectral response acceleration for a 5% damped response spectrum with 2% probability of exceedance within a 50-year period, multiplied by the risk coefficient, CR : defined at short, CRS , and 1-second period, CR1 . The values of the response coefficient can be determined from Figs. 22-18 and 22-19 in ASCE (2017a). For periods less than 0.2 second, CR 5 CRS , and for periods greater than 1.0 s, CR 5 CR1 : For periods in between a linear interpolation between the risk coefficients is allowed. When dealing with tall buildings, where the fundamental period is usually long enough (e.g., T . 1 s), considering CR 5 CR1 is acceptable. b. Spectral ordinate from iterative integration of a site-specific hazard that has collapsed fragility probability density function lognormally distributed. The ordinate of the probabilistic ground motion response spectrum at each period shall achieve a 1% probability of collapse within a 50-year period for a collapse fragility having a 10% probability of collapse at said, and a logarithmic standard deviation of 0.6 (ASCE, 2017a). 2. Deterministic risk-targeted MCER ground motions: at each period, the spectral response acceleration shall be calculated as the 84th-percentile of 5% damped spectral response acceleration in the direction of maximum horizontal response computed at that period. The maximum value of such acceleration calculated for the characteristic earthquakes on all known active faults within the region shall be used. (ASCE, 2017a) Each ordinate shall not be less than the MCER response spectrum as determined for S1 , 0.6, constructed with SS 5 1:5 and S1 5 0:6. Having calculated the site-specific MCER acceleration, SAM , the design spectrum values, SA , can then be calculated as follows: 2 SA 5 SAM 3

(5.14)

Subsequently, the spectral displacement shall be defined as: • • •

SDS as the maximum between SDS determined at 0.2-second period and 90% of SDS determined for any period longer than 0.2-second. SD1 as the maximum between SDS determined 1-second period and double of SDS at 2-second period. SMS and SM1 as 1.5 times SDS and SD1 , respectively.

273

274

CHAPTER 5 Design procedures for tall buildings

•

Design and maximum response displacement values shall not be less than 80% of those determined for S1 , 0.6.

The above-explained site-specific earthquake hazard analysis can be performed with computer software, such as EZFRISK (Risk Engineering, 2012). This is based on the models by Cornell (1973) and McGuire (1976) on probabilistic seismic hazards. This method allows the modeling of the faults, in the vicinity of the building site, as linear sources with fault earthquake activities based on historical and geologic data. Selection and scaling of ground motion records. In the case of nonlinear response procedures, the selection and scaling of appropriate horizontal ground motion acceleration time histories, for MCER, are essential to produce meaningful results (ASCE, 2017a). According to design standards (ASCE, 2017a; PEER, 2017), the following points can be explored: •

• •

•

•

•

At least 11 pairs of ground motions shall be selected for each target spectrum. The chosen events shall have similar characteristics and spectral shape of the target spectrum. In the case of near-fault sites the proportion of ground motion having these effects shall be such that they represent the probability that MCER will have these effects (ASCE, 2017a). Synthetic/simulate ground motions while permitted by ASCE (2017a) shall not be utilized based on design practice. Ground motion modification can be carried out through amplitude-scale or spectral matching (not possible for near-fault site unless the pulse characteristics are maintained after matching) (ASCE, 2017a). The period bounds for scaling or matching are defined as follows (ASCE, 2017a): • Upper-bound period shall be greater or equal to 2 times the largest first mode of the building among the two principal directions (a lower value of 1.5 can be used if justified by dynamic analysis under MCER). • Lower-bound period shall be defined such that it includes a sufficient number of building modes to reach 90% of mass participation in each principle direction. Moreover, it shall not exceed 20% of the lowest first mode among the two principal directions. Instead, in case vertical component is considered, the lower bound period shall not be taken less than 0.1 second or the lowest period (among the two principal directions) which mass participation contribution. Amplitude scaling: for each ground motion pair a maximum-direction spectra shall be constructed. The same factor shall be applied for each ground motion pair such that the average maximum-direction spectra from all ground motion match or exceed the target response spectrum. In any case, it shall not fall below 90% for any period within the selected period range. In case vertical response is considered, each component shall be scaled to envelope the target vertical response spectrum over the period range specified. Spectrum matching: each pair of ground motions shall be modified such that the suite average maximum-direction spectra equal or exceed 110% of the

5.2 Passive Damping Systems

•

target spectrum over the period range specified. For the vertical component the ground motion shall be spectrally matched such as the average of the suite never falls below the target vertical spectrum for the period range specified. Ground motions shall be applied to the support of the model in the building orthogonal direction “such that the average of the component response spectrum for the records applied in each direction is within 10% of the mean of the component response spectra of all the records applied in the period range specified” (ASCE, 2017a). In the design process, the average of the peak responses among a set of ground can be utilized. However, some criteria require to use the maximum of peak response, such as per PEER (2017) in which the peak drift has to be less than 4.5%.

5.2.1.1.3 Step 1.3: Wind demand The design wind load according to ASCE 7-16 (ASCE, 2017a) can be determined based on wind speed that is tabulated (based on risk category, Table 5.7) in the wind hazard map (just for US regions). Except for cases in which regional climate data show unusual wind conditions, the following categories are used in code: • • • •

Risk Risk Risk Risk

category I, Fig. 26.5-1.A (ASCE, 2017a) category II, Fig. 26.5-1.B (ASCE, 2017a) category III, Fig. 26.5-1.C (ASCE, 2017a) category IV, Fig. 26.5-1.D (ASCE, 2017a)

Two procedures are commonly utilized for determining wind loads for static analysis: directional and wind-tunnel procedures (ASCE, 2017a). The latter procedure can be applied to all types of buildings while the former has several limitations (i.e., applicable only to regular-shaped buildings, no effects on building due to cross-wind loading, vortex shedding and instability due to galloping or flutter, and building site location with no channeling effect or buffeting in the wake of upwind obstructions). These requirements show that dynamically sensitive buildings (i.e., tall buildings) might require wind-tunnel testing. Directional wind procedure. This procedure can be applied to all types of buildings that are regular shaped and are not subject to a cross-wind loading (ASCE, 2017a). In all other cases, the designer shall use the wind-tunnel procedure, as explained in the next substep. The code provides two different procedures depending on the height: above or below the limit of 48.8 m (160 ft). In this book, only the procedure for structures with height above 48.8 m will be reviewed. The procedure proposed by the ASCE 7-16 (ASCE, 2017a) considers the application of wind in each direction as an independent action: windward, leewar, and side walls. There are eight different steps in the procedure, as shown in Fig. 5.10. 1. Have been already discussed in Section 5.2.1.1.1. 2. Have been already discussed in Section 5.2.1.1.3.

275

276

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.10 ASCE (2017a) directional wind procedure steps.

3. The different wind-load parameters to be calculated are the following: • Wind directionality factor, Kd : determined from Table 5.12. • Exposure category: for each wind direction, shall be based on the surface roughness and exposure categories. These need to be determined for the two upwind section extending 45 degrees on each side. For further details, refer to Section 26.7 of ASCE 7-16 (ASCE, 2017a). • The topographic factor, K zt : it takes into account the changes in the topography of the proximity of the structure under analysis. For further details, refer to Section 26.8 of ASCE 7-16 (ASCE, 2017a). • Gust effect factor, Gef : it takes into account the wind turbulence
structure interaction, as well as the dynamic amplification due to building flexibility. This factor is computed differently for rigid (,1 Hz) and flexible (.1 Hz) buildings. For rigid buildings, it is permitted to be simply taken as 0.85 or computed as follows:  1 1 1:7gQ Iz Q Gef 5 0:925 1 1 1:7gv Iz

(5.15)

5.2 Passive Damping Systems

Table 5.12 Wind Directionality Factor (ASCE, 2017a) Directionality Factor, Kd

Structure Type Buildings Main wind force-resisting system Components cladding Arched roofs

0.85 0.85 0.85

Chimneys, Tanks, and Similar Structures Square Hexagonal Round

0.90 0.95 1.00

Solid freestanding walls and solid freestanding and attached sign

0.85

Open signs and lattice framework

0.85



10 1=6 z vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi u 1 Q5u  0:63 u t 1 1 0:63 L2L1H z Iz 5 gc

 A z Lz 5 gl 10

(5.16)

(5.17)

(5.18)

where Iz is the intensity turbulence at the equivalent height of the building, z is defined as 0:6H (but not less than zmin , Table 5.13); Q is the background response; the constants gQ and gv shall be taken as 3.4; the constants gc , gl , and A are shown in Table 5.13; L2 is the building horizontal dimension perpendicular to the wind direction; and H is the mean roof height. Instead, for flexible buildings the gust factor, Gef , can be computed as follows: qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi1 0 1 1 1:7Iz g2Q Q2 1 g2R R2G A Gf 5 0:925@ 1 1 1:7gv Iz

(5.19)

The constants gQ and gv shall be taken as 3.4 and gR as follows: gR 5

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 0:577 2 lnð3600f1 Þ 1 pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 lnð3600f1 Þ

(5.20)

277

Table 5.13 Gust Effect Factor Constants (ASCE, 2017a) Exposure

α

zg (m)

a^

b^

α

b

gc

gl (m)

A

zmin (m)a

B C D

7.0 9.5 11.5

365.76 274.32 213.36

1/7.0 1/9.5 1/11.5

0.84 1.00 1.07

1/4.0 1/6.5 1/9.0

0.45 0.65 0.80

0.30 0.20 0.15

97.54 152.4 198.12

1/3.0 1/5.0 1/8.0

9.14 4.57 2.13

The equivalent height z is equal to the greater of 0:6H and zmin and for building with H # zmin is equal to zmin .

a

5.2 Passive Damping Systems

where f1 is the building fundamental frequency and RG is the resonant factor: sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 RG 5 Rn Rh RB ð0:53 1 0:047RL Þ ζT Rn 5

7:47N1

(5.22)

ð1110:3N1 Þ5=3 n1 Lz Vz

(5.23)

 1 1  2 1 2 e22η for η . 0 η 2η2

(5.24)

N1 5 Rl 5

(5.21)

Rl 5 1 for η 5 0

(5.25)

where the subscript l in Eqs. (5.24) and (5.25) shall be taken as H, L2 , and L1 (building horizontal dimension perpendicular and parallel to wind direction); ζ T is the total damping ratio; and η is a constant equal to: Rl 5 RH -η 5 4:6f1 H=V z

(5.26)

Rl 5 RL2 -η 5 4:6f1 L2 =V z

(5.27)

Rl 5 RL1 -η 5 15:4f1 L1 =V z

(5.28)

where V z is the mean hourly wind speed at height z:  α z V Vz 5b 10

• •

(5.29)

where b, α, and z are defined in Table 5.13 and V is the basic wind speed. Enclosure classification: categorized buildings as enclosed, partially enclosed and open. Internal pressure coefficient, GC pi : based on the enclosure classification, as shown in Table 5.14. For large volume building, partially enclosed, a reduction factor can be utilized as: 0

1

1 B C RGCpi 5 0:5@1 1 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiA Vvi 1 1 22;800Aog

(5.30)

where Aog is the opening total area in the building envelope (walls and roof in ft2) and Vvi is the unpartitioned internal volume (in ft3). 4. The velocity pressure exposure coefficients, Kz or Kh , are determined from the following equation (Table 5.15):  2=α 4:57m # z # zg -Kz 5 2:01 z=zg  2=α z , 4:57m-Kz 5 2:01 15=zg

(5.31)

where zg and α are tabulated in Table 5.13. Linear interpolation for intermediate values of z is allowed (ASCE, 2017a).

279

280

CHAPTER 5 Design procedures for tall buildings

Table 5.14 Internal Pressure Coefficient (ASCE, 2017a) Enclosure Classification

GCpi

Open buildings Partially enclosed buildings

0.00 1 0.55 20.55 1 0.18 20.18

Enclosed buildings

Note: Plus and minus signs indicate pressure acting toward and away from the surface, respectively. Two cases shall be considered: one with positive value of GCpi applied to all surfaces and one with negative values.

Table 5.15 Velocity Pressure Exposure Coefficients (ASCE, 2017a) Exposure Height Above Ground Level (m)

B

C

D

0
4.6 6.1 7.6 9.1 12.2 15.2 18 21.3 24.4 27.4 30.5 36.6 42.7 48.8 54.9 61.0 76.2 91.4 106.7 121.9 137.2 152.4

0.57 0.62 0.66 0.70 0.76 0.81 0.85 0.89 0.93 0.96 0.99 1.04 1.09 1.13 1.17 1.20 1.28 1.35 1.41 1.47 1.52 1.56

0.85 0.90 0.94 0.98 1.04 1.09 1.13 1.17 1.21 1.24 1.26 1.31 1.36 1.39 1.43 1.46 1.53 1.59 1.64 1.69 1.73 1.77

1.03 1.08 1.12 1.16 1.22 1.27 1.31 1.34 1.38 1.40 1.43 1.48 1.52 1.55 1.58 1.61 1.68 1.73 1.78 1.82 1.86 1.89

5.2 Passive Damping Systems

5. The velocity pressure, qz, can be determined at height z and at the mean roof height H, qH , as follows: qz 5 0:613Kz Kzt Kd V 2

(5.32)

where V is the basic wind speed in m/s, Kz is the velocity pressure coefficient, Kzt is the topographic factor, and Kd is the wind directionality factor, as previously defined. In case z 5 H, the velocity pressure, qH , is computed at the roof height. 6. The external pressure coefficient, Cp and CN , can be determined based on the charts provided in Section 27.3 of ASCE 7-16 (ASCE, 2017a). 7. The wind pressure, p, can be determined on each building surface depending on the type of building: • Enclosed and partially enclosed rigid buildings: p 5 qGCp 2 qi GCpi

(5.33)

•

where q 5 qz for windward walls at height z; q 5 qH for leeward and side walls, and roofs at height H; qi 5 qH for windward, leeward, and side walls, and roofs at height H of enclosed buildings and for negative internal pressure for partially enclosed buildings; and qi 5 qz internal pressure for partially enclosed buildings with z evaluated at the highest opening influencing the positive internal pressure. Conservatively, qi 5 qH can be taken. Enclosed and partially enclosed flexible buildings:

•

Open buildings:

p 5 qGf Cp 2 qi GCpi

(5.34)

p 5 GCN

(5.35)

where CN is net pressure coefficient as determined in step 6. The wind pressure as calculated above shall not be less than the following requirements: • For enclosed and partially enclosed buildings: 0.77 kN/m2 for walls and 0.38 kN/m2 for roofs. • For open buildings: 0.77 kN/m2. 8. The action on each surface shall be applied concurrently on windward walls, leeward walls, and roofs, as shown for the load cases of Fig. 5.11. The figure shows how the influence of eccentricity is a crucial element and the values provided refer to rigid buildings calculated for each main direction (ex and ey ). Instead, flexible buildings shall be computed as follows, for each main direction: e5

qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi  2 gQ QeQ 1 ðgR RG eR Þ2 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 gQ Q 1 ðgR RG Þ2 1 1 1:7Iz

eq 1 1:7Iz

(5.36)

where eQ is the eccentricity, e, for rigid structures (Fig. 5.11), and eR is the distance between the elastic shear center and the center of mass of each floor. Wind-tunnel procedure. In case the previously described directional procedure is not applicable, wind-tunnel tests are required to determine the wind pressure on

281

282

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.11 Design wind load cases. Adapted from ASCE, 2017a. ASCE7-16: Minimum Design Loads for Buildings and Other Structures. American Society of Civil Engineers, Reston, VA.

the building, and also for testing the efficiency of dynamic modification system (see Chapter 7 for more details). The conditions, for the tunnel testing protocol, required by ASCE 7-16 (ASCE, 2017a) and ASCE 49 (ASCE, 2012) are as follows: • •

•

Modeling of wind-speed variation with height is necessary (i.e., atmospheric boundary layer). Accurate scale modeling is essential to model atmospheric turbulence and the geometric characteristics of the surrounding environment, as well as the building under study. The model’s projected area is less than 8% of the total test section area, unless a correction is considered for blockage.

5.2 Passive Damping Systems

• • • •

Accounting for longitudinal pressure gradient is necessary. Must attempt to minimize the Reynolds number of pressures and forces. Response parameters are estimated, based on the required measurements of the tests. The structural model of the building under study shall take into account a correct distribution of mass, stiffness, and damping.

The load effects on buildings shall have the same recurrence interval as that for analytical methods recognized in the existing literature. These methods would allow the combination of the wind-tunnel test data with meteorological data or a probabilistic model.

5.2.1.1.4 Step 1.4: Load combinations The effects of gravity loads and seismic/wind forces can be combined as follows (ASCE; 2017a): •

Basic load combinations for strength design Earthquake: 

ð1:2 1 0:2SDS ÞDL 1 LL 1 EL 1 0:2SL ð0:9 2 0:2SDS ÞDL 1 EL 1 1:6HL

(5.37)

Wind: 

•

1:2DL 1 LL 1 WL 1 0:5SL 0:9DL 1 WL

(5.38)

Basic load combinations for allowable stress design Earthquake: 8
> > > > R > > T@ A > > < Ie

8 9 0:044SDS Ie $ 0:01 > > > > > > > > < = 0:5SD1 I !e for S1 $ 0:6 # Cs 5 SDS # SD1 TL > > > R > > > 0 1 for T1 . TL > > > Ie > IRe > > : > ; > R > > 2 > > T @I A :

(5.45)

e

The important factor, Ie , is defined in Table 5.7 and the response modification coefficient, R, can be determined using Table 5.16; SD1 , SDS , and TL are determined previously in Step 1.2 (Section 5.2.1.1.2); T1 is the fundamental period of the building; and Bζ is the damping reduction factor. Note that in Eq. (5.44) the limit, 0:75Cs W, is not applicable in case less than two dampers per main direction in each story are provided and if horizontal and/ or vertical irregularities are present (ASCE, 2017a).

291

292

CHAPTER 5 Design procedures for tall buildings

Determine the parameters of the lateral force
resisting system. ASCE 7-16 (ASCE, 2017a) allows the designer to take into account the inelastic behavior of the structural system and provides simplified formulas to estimate the relative hysteretic damping. It is important to note that buildings equipped with dynamic modification devices are usually designed to remain elastic for all the levels of excitation. In case inelastic deformations are considered, in the design process, it should be proven that these are not detrimental for the behavior of the damping system (ASCE, 2017a). In any case the parameters to take into consideration, when considering inelastic deformations, are the following: •

•

Effective ductility demands μD and μM : The effective ductility demand of the seismic force-resisting system of the building is defined as the ratio between the fundamental mode displacement, D1 ; and the displacement at the effective yield point of the system, Dy , both determined at the center of rigidity of the roof level in the direction of interest (ASCE, 2017a; NEHRP, 2015). Depending if D1 is estimated for the DE or the MCER, ground motions, the effective ductility is specified using symbol μD or μM , respectively. For the RSA of the damped structural system, the effective ductility demands are utilized for taking the inelastic (nonlinear) structural effects on the responses, as well as for updating the hysteretic and total damping contributions into account. As an initial phase, the designer may assume values of the effective ductility demands on the structure due to the DE (μD ) and MCER (μM ). These parameters may depend on the structural system and material type at hand. As discussed earlier, it is a general assumption to assume μD 5 μM 5 1, since the elastic behavior of the damped structure is desired. Effective fundamental mode period: In current building codes (ASCE, 2017a; NEHRP, 2015) nonlinearities are taken into account the global building behavior through the definition of the effective fundamental period. Depending upon the earthquake level (i.e., DE or MCER ground motions), this period can be explicitly computed by considering the postyield force
deflection characteristics of the structural system or can be obtained as follows: pffiffiffiffiffiffi T1D 5 T1 μD pffiffiffiffiffiffiffi T1M 5 T1 μM

•

(5.46) (5.47)

Modal seismic base shear and modal effective seismic weight: The modal base shear, Vm , of the mth mode in the desired direction can be readily calculated as: Vm 5 Csm Wm

(5.48)

where Csm is the modal seismic response coefficient (computed as shown below) and Wm is the modal effective seismic weight of the mth mode of structure given by the following:

5.2 Passive Damping Systems Pn

2 wi φim Wm 5 Pn  2 i51 wi φim i51

•

(5.49)

where φim is the mth modal displacement amplitude at the ith level of the structure (Section 5.2.1.3) in the direction of interest, normalized to unity at the roof level, and wi is the portion of the total effective seismic weight W, assigned to level i. Modal seismic response coefficients: these coefficients are computed for the fundamental and higher modes as follows: For fundamental modes: 

R SDS for T1D , TS Cd Ω0 Bζ1D

(5.50)

R SD1 for T1D $ TS Cd T1D Ω0 Bζ1D

(5.51)

Cs1 5  Cs1 5

For higher modes (m . 1): 

Csm 5

R SDS for Tm , TS Cd Ω0 BζmD

R SD1 for Tm $ TS Csm 5 Cd Tm Ω0 BζmD

(5.52)



•

(5.53)

Here, response modification coefficient R, deflection amplification factor Cd , and overstrength factor Ω0 were determined using Table 5.16; Tm is the period of the mth mode of vibration in the direction of interest (Step 3, Section 5.2.1.3); SD1 and SDS are determined in Step 1 (Section 5.2.1.1); and BζmD is the mth mode numerical damping reduction factor (see Table 5.18) for mth mode supplemental damping and Tm $ T0 5 0:2 SSD1 . As previously stated it DS is common practice to assume the same value of the supplemental damping for all modes. Total seismic base shear V and check its limit: ASCE (2017a) standard requires that the seismic base shear and its minimum value should be computed for a given direction of the interest. Provided that closely spaced modes do not have significant cross-correlation, the total base shear can be calculated with the square root of sum of the squares (SRSS) by: vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi u Nm uX V 5t ðVm Þ2 $ Vmin

(5.54)

m51

where V min is calculated in Eq. (5.44). As an alternative, a CQC can be adopted to compute the base shear (refer to Chopra (2006) for further details).

293

294

CHAPTER 5 Design procedures for tall buildings

•

Note: For a low-rise building with less number for mode shape, this procedure is practical. Instead, for tall building for which the analysis could include more than 100 mode shape, this method may be impractical. Design lateral force in modal form (Fim ): To determine the forces in the structural elements the design base shear needs to be distributed along the building height. ASCE 7-16 (ASCE, 2017a) proposed the calculation of the design lateral force, Fim , at ith level due to the mth mode in a given direction as follows: Fim 5 wi φim

Γm Vm Wm

(5.55)

where Γm is the modal participation factor associated for the mth mode of vibration in the direction under consideration that can be estimated as follows: Wm i51 wi φim

Γm 5 Pn

(5.56)

•

Design force in structural elements: Design forces in the structural elements can be determined using the SRSS or CQC of modal design forces (Chopra, 2006). To this end, the design lateral forces from the previous equation can be utilized in commercial software with static analysis. The alternative is to employ RSA directly, in a structural software with the parameters previously determined. Determine the parameters of distributed damping system. Based on the design and MCER seismic demands, the damper parameters can be separately determined with the use of the subsequent steps (ASCE, 2017a). •

DE roof displacement in modal forms (DmD m 5 1; 2; . . . ; n): For fundamental modes: D1D 5

2 g SDS T1D g SDS T12 Γ1 $ 2 Γ1 for T1D , TS 2 BζT1D BζT1E 4π 4π

(5.57)

D1D 5

g SD1 T1D g SD1 T1 Γ1 $ 2 Γ1 for T1D $ TS 4π2 4π BζT1D BζT1E

(5.58)

For higher modes (m . 1): DmD 5

•

g SD1 Tm g SDs Tm2 Γm # 2 Γm 2 4π 4π BζTmD BζTmD

(5.59)

where BζT1D and BζTmD are the total damping reduction factors for the fundamental and mth mode, respectively, at design displacement, and BζT1E is the total damping reduction factor for the fundamental mode at initial displacement (Table 5.18). DE floor deflection in modal δi;mD and total δiD forms: δi;mD 5DmD φim

(5.60)

5.2 Passive Damping Systems

vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi u N uX  2 δiD 5 t δi;mD

(5.61)

m51

•

DE story drift in modal Δi;mD and total ΔiD forms: The story drift, Δi;mD , of mth mode and total drift, ΔiD , both associated with ith story can be computed as: Δi;mD 5 δi;mD 2 δi21;mD vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi u Nm uX  2 Δi;mD ΔiD 5 t

(5.62) (5.63)

m51

•

DE story velocity in modal rimD and total forms riD : For fundamental modes: ri;1D 5 2π

Δi;1D T1D

(5.64)

ri;mD 5 2π

Δi;mD Tm

(5.65)

For higher modes (m . 1):

The total story velocity associated with ith story is: vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi u Nm uX  2 riD 5 t ri;mD

(5.66)

m51

•

MCER roof displacement in modal forms (DmM form 5 1; 2; . . . ; n): For fundamental mode: D1M 5

2 g SMS T1M g SMS T12 Γ1 $ 2 Γ1 for T1M , TS 2 BζT1M BζT1E 4π 4π

(5.67)

D1M 5

g SM1 T1M g SM1 T1 Γ1 $ 2 Γ1 for T1M $ TS 4π2 4π BζT1M BζT1E

(5.68)

For higher modes (m . 1): DmM 5

g SM1 Tm g SMs Tm2 Γm # 2 Γm 2 BζTmM BζTmM 4π 4π

(5.69)

where BζT1M and BζTmM are the total damping reduction factors for the fundamental and mth mode, respectively, at maximum displacement (Table 5.18). In the subsequent steps, the structure inelastic behavior is taken into account while calculating the total damping (ASCE, 2017a; NEHRP, 2009, 2015). Using these steps causes iterations, and designers should repeat the calculations of this section until convergence.

295

296

CHAPTER 5 Design procedures for tall buildings

Determine ductility demands. The maximum ductility associated with the structural system can be calculated as: " μ max 5 0:5

R Ω 0 Ie

μ max 5

#

2 11

for T1D # TS

R for T1 $ TS Ω 0 Ie

(5.70)

(5.71)

where Ie can be chosen from Table 5.7. This maximum value shall not be exceeded the ductility for DE, μ D , and MCER, μ M , levels, computed as follows: μD 5

4π2 RD1D $1 gCd Ω0 Γ1 CS1 T12

(5.72)

μM 5

4π2 RD1M $1 gCd Ω0 Γ1 CS1 T12

(5.73)

Note that, alternatively, it is permitted (ASCE, 2017a; NEHRP, 2015) to use nonlinear static (pushover) analysis, to develop a force
displacement curve of the system and to use this curve instead of the above equations in order to calculate the ductility demands. Determine hysteretic damping ratios ζ hD and ζ hM . In this step, with the use of inherent damping ratio, ζ i , specified in Step 3 (Section 5.2.1.3), the hysteretic damping ratio can be calculated, ζ hD and ζ hM , which are the damping contribution postyield hysteretic behavior for the DE and MCER levels, respectively, at effective ductility demands μD and μM , determined by: 8 0 > >   > > ζ hD 5 qH 0:64 2 ζ i @1 2 > > > < 0 > > >   > > ζ hM 5 qH 0:64 2 ζ i @1 2 > > :

1 1A μD 1 1 A μM

(5.74)

where 0:5 # qH 5 0:67

TS #1 T1

(5.75)

Determine the modal effective damping for design and maximum ground motions. This step gives the designer the opportunity to update the total effective damping by accounting for the ductility demands (μ D and μ M ) under DE and MCER ground motions, respectively. To this end, considering the hysteretic damping ratio (ζ h) and by including μ D and μ M , one can calculate the updated total damping ratio using Eqs. (5.76) and (5.77), respectively. Afterward, the

5.2 Passive Damping Systems

designer should refer to the above paragraphs to modify the structural and damping responses. In case nonlinear structural behavior is not considered, this step (as well the previous ones) is neglected. 

pffiffiffiffiffiffi ζ T1D 5 ζ i 1 ζ d1 μ D 1 ζ hD ζ TmD 5 ζ i 1 ζ dm 1 ζ hD for m . 1  pffiffiffiffiffiffiffi ζ T1M 5 ζ i 1 ζ d1 μ M 1 ζ hM ζ TmM 5 ζ i 1 ζ dm 1 ζ hM for m . 1

(5.76) (5.77)

Here, ζ hD and ζ hM are determined previously. The supplemental damping ratio, ζ d1 , at the fundamental mode is estimated in Step 5 (Section 5.2.1.5). It should be noted that the damping ratio, ζ dm , in higher modes may be assumed identical to that of the fundamental mode (i.e., ζ dm 5 ζ d1 ), or can be alternatively calculated using equations from other sources (see Lago (2011) for more information).

5.2.1.4.2 Step 4.2: Nonlinear procedure ASCE 7-16 (ASCE, 2017a) allows the utilization of nonlinear procedures for all types of buildings. Two different types of analyses are allowed: 1. Nonlinear response-history procedure 2. Nonlinear static procedure In both procedures, the structural and dynamic modification element’s nonlinear properties shall be explicitly modeled as verified by testing (ASCE, 2017a). The damper model shall also take into account the frequency and amplitude dependencies. However, this is not required when the properties vary during the loading history. In this case, an upper and a lower bound of the device properties can be used and the envelope of the dynamic response can be utilized (refer to Step 8 (Section 5.2.1.8.3) for more details). Nonlinear response-history procedure. This procedure considers the 3D numerical modeling of the whole building structural system, taking into account its intrinsic nonlinear hysteretic behavior, except for the elements of force-resisting system with ductility less than 1.5 that could be modeled as linear elements (ASCE, 2017a) (Note: careful consideration about modeling linear elements should be taken care since this is considered valid for systems with focused ductile behavior.) The nonlinear behavior should be consistent with test data and shall represent the overall cyclic behavior (i.e., yielding, strength and stiffness degradation, and pinching). The model shall be assumed as fixed based unless a realistic SSI model is used. Diaphragm flexibility shall be taken into account the case they are not rigid, compared to the vertical elements of the structural system. The numerical model should consider the actual spatial distribution of the mass throughout the structure. Nonlinear static procedure. Nonlinear static “pushover” procedure is used in addition to response-spectrum and ELF procedures in the case S1 . 0:6 to confirm

297

298

CHAPTER 5 Design procedures for tall buildings

the peak response (NEHRP, 2015). This procedure shall utilize a numerical model, similar to that of the nonlinear response-history procedure. The result of the analysis shall be a force
displacement curve that can be utilized for determining the ductility demand, as shown in Section 5.2.1.4.1.

5.2.1.5 Step 5: Select total target damping A trial supplemental equivalent damping ratio, ζ d , should be chosen by the designer. As shown in Chapter 3 (Section 3.2), this value is just one part of the building total damping that is made of this additional contribution: intrinsic, aerodynamic, and hysteretic. The intrinsic damping estimation is not trivial and ASCE 7-16 (ASCE, 2017a) specifies that it should be not greater than 3% for all the modes (detailed discussion on the different standard requirements was done in Chapter 3 (Section 3.2.2)). The hysteretic damping is commonly neglected (i.e., ζ h 5 0) in practice for buildings with dynamic modification devices, since it could be detrimental for the behavior of the dissipating system. In any case, a more detailed discussion on how to calculate structural hysteretic damping is given in Section 5.2.1.4.1. Similarly, the aerodynamic damping is neglected but it could become relevant with the next generation of slender supertall buildings. The amount of total damping is usually decided based on a combination of code requirements, economical, and practical aspects. Looking at code requirements, US code (NEHRP, 2015) allows a maximum total damping of 35%. Instead, the Chinese code (GB50011, 2010) recommends a total damping of less than 25%. In literature, other examples are found: Lee and Taylor (2001) suggest a value between 15% and 25%, while Zhou et al. (2012) suggested to use 15%. Note that this estimated damping value refers to the building fundamental mode of vibration and it is important to specify it also for higher modes. This is especially relevant for tall buildings that, in most cases, do not have single dominant modes. It is common practice to assume that the inherent damping is constant through all the modes, but this could be not conservative enough for the damping of the dissipation and hysteretic damping (see Lago (2011) for more information on different procedures available for estimating the higher mode damping). Starting from these recommendations, a designer can select a total damping value based on the required reduction of the building demand obtained from the target spectrum and the estimated predominant modal periods (Step 3, Section 5.2.1.3).

5.2.1.6 Step 6: Damper type, configuration, and distribution In this step, designers should choose the dynamic modification technology to utilize and this is obtained from the following three tasks: type of dampers, installation configuration of dampers, and distribution of devices over the building. 1. Selection of energy-dissipating devices: In general, distributed dampers can be classified into two categories: (1) velocity-dependent and (2) displacement-

5.2 Passive Damping Systems

dependent devices (for more details refer to Chapter 4 (Section 4.1.1)). The designer should select the appropriate type of distributed dampers with regard to several aspects, including: a. Type and form of tall building structural system: Depending on the type and form of the structure (e.g., frame, wall systems, combined systems) the type of damper may vary. For example, for coupled shear wall buildings friction dampers (Chung et al., 2009) and viscoelastic devices (Christopuolus and Montgomery, 2013) may be more suitable as coupling elements. b. Damper economics: In general, displacement-dependent devices are less expensive than velocity-dependent ones (e.g., Liang et al. (2012)). Also, the cost related to the availability of devices, depending on the site location, may be important. Moreover, displacement-dependent devices may need to be replaced after an event (see Chapter 3 (Sections 3.6) for further details about the economic aspects of damping devices). c. Target performance: Depending on the structural performance to be achieved, the selected type of devices may differ. For instance, velocitydependent dampers might slightly increase the base shear at the foundation level when the structure becomes inelastic, while displacement-based devices exhibit less of these problems (Symans et al., 2008). Furthermore, the use of viscous dampers more effectively reduces vibrations (Liang et al., 2012). d. Excitation force: Depending on which external lateral load (e.g., wind or earthquake) is exciting the tall buildings, the type of dampers applied may be different. For example, velocity-dependent dampers are usually preferred for wind-dominant excited buildings, rather than displacementbased ones. Additionally, displacement-dependent dampers, especially friction dampers, are more suitable for buildings located near high-seismic zones with long-period ground motion histories (Liang et al., 2012). Further details and guidelines about the selection of damper type are provided in Chapter 4 (Section 4.4). 2. Damper geometrical configuration: After selecting the type of energydissipating devices, the most suitable device configuration can be chosen by the designer, mainly depending on the type of lateral load
resisting structural system. The most common damper configurations are represented in Chapter 4 (Section 4.1.1). In case the tall building’s lateral system consists of a frame-like (shear-type) structure, the configurations illustrated in Fig. 4.2 may be adopted. Instead, if coupled walls are used, the installations shown in Fig. 4.16 or 4.19 could be more appropriate. In addition, diagonal or Chevron configurations could not be the preferred solution due to architectural constraints (see Chapter 6 (Section 6.1) for further details). Another reason could be the presence of stiff coupling beams and shear wall systems with relative small interstory drift velocities that might not be sufficient to activate diagonal or Chevron damper configurations (Lavan, 2012).

299

300

CHAPTER 5 Design procedures for tall buildings

3. Damper distribution: Given the type and damper geometrical configuration, the device distribution along the plan and height of the building can be determined by the designer, based on the following considerations: a. Plan distribution. Concerning the building plan, the distribution should be symmetrical (ASCE, 2017a) to the center of stiffness in such a way that no torsional effects could develop due to damper forces under seismic or wind loading. To control torsional effects, it is recommended (NEHRP, 2015) that at least four dampers are considered in each story and along each main direction of building, with at least two devices (per direction) on each side of the center of stiffness of any story. If these minimum conditions are not satisfied, all dampers must be capable of sustaining 1.3 times the maximum displacements obtained under MCER; moreover, velocity-dependent dampers must sustain displacements and forces associated with 1.3 times the maximum velocity under MCER. No additional requirements are given for wind design. Depending on the most critical direction under external excitations, the damping system may be assigned to one direction of the building plan or to more directions; the desirable direction to be equipped with dampers in existing buildings may sometimes depend upon the possible location for the device placement within the structural systems in that direction. For a case study example of this type, see San Diego Courthouse case study, in Chapter 8 (Section 8.1.6). b. Vertical distribution. Regarding the distribution in height, in general, these general aspects should be considered: interstory drifts and velocities profile, damper mechanical properties distribution in height, and the number of devices assigned in each story. Building codes usually provide general recommendations on how dampers distribute vertically. American standards (e.g., FEMA (2000d)) suggest that damper should be arranged evenly along the height of the building. In China, GB50011 (2010) suggests that energy dissipation devices can be positioned in those stories where the drift is larger. Its number and distribution should be decided reasonably by means of comprehensive analysis, so that it is beneficial to improve the energy dissipation capacity of the whole structure. However, GB50011 (2010) provides the formula for calculating the effective supplemental damping ratio, which is applicable only when the dampers are arranged evenly along the height of the building. Based on these considerations, the traditional arrangement scheme for dampers is to place them in the story with larger drifts or velocities, and at the same time, try to arrange them evenly along the height of the building. However, in practical high-rise projects, the two aspects above are hardly achievable. Current super tall buildings are mainly composed of framecore tube structures with outriggers at several floors along the height. In this case, dampers could be placed only at outriggers stories that usually

5.2 Passive Damping Systems

have the smallest story drift and velocity (see Section 5.6.2 for further distribution optimization criteria). In literature the major several vertical distribution methods can be found and the most common are the following: uniform, based on damper properties (Whittle et al., 2012) and inversely proportional distribution (Takewaki, 2009). The first approach has been successfully applied in several different buildings: uniform distribution of viscoelastic dampers within the original World Trade Center in New York City and the Santa Clara County Building in San Jose, California (Soong and Dargush, 1997, and see Chapter 8 for more case studies). However, this distribution method may not be the most effective one because it does not take into consideration the different demand at each floor. To overcome this, a partial distribution of identical dampers around the maximum shear deformation along the height could be utilized (Lago, 2011; Christopoulos and Montogomery, 2013). The second basic approach is based on distributing dampers based on their mechanical properties. The major methods are stiffness- and massproportional distribution. The former considers the distribution of the dampers with the relative constant, c, proportional to the ith story lateral stiffness, ki . The larger the story stiffness, ki , the greater the damping coefficient, ci . The stiffness can be estimated simply from applying a distributed static lateral load and the ratio of interstory shear force to interstory displacement at each level gives the lateral story stiffness for a typical structural system (Whittle et al., 2012). The other option can be based on mass-proportional distribution in which the damper constant is proportional to the mass. In this case, dampers are placed between the floor and a fix point (usually the ground), but this is quite impractical for tall buildings (Trombetti and Silvestri, 2006). The third approach is based on the work by Takewaki (2009) and Takewaki and Fujita (2009). They emphasized that the distribution of passive energy-dissipating devices is not efficient in upper stories of tall buildings. Moreover, Takewaki (2009) mentioned that the dampers are more effective for shear deformation; thus, the linear distribution of dampers, consistent with dominance of such a deformation along tall buildings, is proposed. The proposed simplified strategy (Takewaki, 2009) consists in multiplying a constant damping coefficient by a series of factors, defined by the user, to obtain an inverse-linear distribution (see Fig. 5.14). Only two coefficients, less than one, are needed: υ1 assigned to the first story and υN associated with the topmost story. In addition, to the above-mentioned distributions, some other advanced methods can be employed to get the optimal distribution of dampers along the building height. These include sequential search algorithm (SSA)-mode application (Lopez-Garcia, 2001) and analysis/redesign approach (Lavan and Levy,

301

302

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.14 Fully linear distribution of damping along tall building height.

2006) (further details about more advanced techniques are described in detail in Section 5.6).

5.2.1.7 Step 7: Damping preliminary design In this step, the preliminary design of the major type of distributed damping devices (viscous, viscoelastic, and hysteretic/friction) is reviewed and discussed in detail.

5.2.1.7.1 Viscous dampers’ properties estimation Generally speaking, the main parameters to define in order to design a viscous damper are velocity exponent, α, damping coefficient, c, and damper power, ED . Velocity exponent. In case nonlinear viscous dampers are chosen, the velocity exponent, α, should be preliminarily selected by the designer. A viscous damper catalogue is commonly provided by manufacturers for several specified values of α (Zhou et al., 2012). Generally speaking, the suitable α value is smaller for the seismic design than wind-resistant design (Taylor, 2010). As discussed in Chapter 4 (Section 4.1.1.1.1), the main advantage of utilizing linear damper (i.e., α 5 1:0) is that the damping force has 90 degrees phase difference with the elastic restoring force of structure. For nonlinear viscous dampers, the phase difference diminishes as the velocity exponent gets smaller (see Fig. 5.15). Damping coefficient. The damping coefficient can be estimated knowing the velocity exponent, α, of viscous dampers and the supplemental target damping ratio, ζ d , chosen by the designer in Step 5 (Section 5.2.1.5). Additionally, to determine the damping coefficient, the type of damping distribution along the building height needs to be chosen. Hence, the designer can choose between different damping distributions, previously described in Step 6

5.2 Passive Damping Systems

FIGURE 5.15 Constitutive relation curves of different damping indeces.

(Section 5.2.1.6): uniform (Liang et al., 2012; Palermo et al., 2016), stiffnessproportional (ASCE, 2017a), and inversely linear (Takewaki, 2009). In the following, each distribution method is briefly reviewed. •

Uniform damping distribution. In case viscous dampers are uniformly distributed, Liang et al. (2012) have proposed a preliminary design method for estimating the damping coefficient according to different installation arrangements. Considering a nonlinear viscous damper with constant, Ci , associated with the ith story (if more than one damper per stories this is the total required damper constant per story), and velocity exponent, α, the generalized expression is developed as follows: Ci 5

12α 4πM1  0:8u_ di;max ζd 2 NT1 ðGah Þi

(5.78)

Here, N is the number of stories that are equipped with dampers; M1 is the fundamental mode effective mass, obtained from modal analysis performed in Step 3 (Section 5.2.1.3); u_ di;max is the damper maximum velocity at the ith floor; ðGah Þi is the geometrical magnification factor that relates to the ith floor to the (i 2 1)th floor (see some common relations of ðGah Þi , as illustrated in Chapter 4 (Fig. 4.2), according to various damper configurations and summarized in Table 5.19)). The maximum interstory velocity, u_ di;max , can be obtained from commercial software results or from empirical relations, as proposed by Palermo et al. (2016) for shear-type buildings. Alternatively, in response to spectrum procedure the maximum of interstory velocity can be calculated from Step 4 (Section 5.2.1.4.1). For preliminary design, Zhou et al. (2012) suggest that a velocity of 200
250 mm/s is usually appropriate for building viscous dampers. Moreover, Sinclair (2006) took a design velocity of 254 mm/s for viscous dampers.

303

304

CHAPTER 5 Design procedures for tall buildings

Table 5.19 Horizontal Magnification Factors for Various Damper Installation Schemes Damper Installation Type Diagonal Intradiagonal Scissor Chevron Bottom toggle

Geometrical Magnification Factor ðGah Þi   cos θj  a cos θj Hjj   sin θj1   tan θj2 1

  cos θj1   cos θj1 1 θj2

  sin θj2   cos θj1 1 θj2     aj :cos θj1   2 cos θj2 cos θj1 1 θj2 bj aj   sin θj1 1

Up-Toggle Modified Toggle Modified Chevron

In alternative to Eq. (5.78) a simplified relation (Silvestri et al., 2010; Palermo et al., 2016) can be utilized as follows: Ci 5

 12α N 11 2πm  0:8:u_ di;max ζd 2 nd T1 ðGah Þi

(5.79)

where nd is the total number of identical viscous dampers placed at each story; N is the number of building stories; and m is the total mass of the building. In the case of buildings with significant flexural deformation (e.g., tall buildings with wall systems) the vertical deformation could be comparable to the horizontal one (as already explained in Chapter 4 (Section 4.1.1.1)). For this reason, Hwang et al. (2008) propose an update of Eq. (5.78) to take into account the vertical deformations for linear viscous dampers with uniform distribution along the height:  2 m i φh i Ci 5 P h     i2 ζ d T1 j ðGah Þj φh rj 2 ðGav Þj φv rj 4π

P

i

(5.80)

where ðGah Þj and ðGav Þj are magnification factors in the horizontal and vertical directions, (see Table 5.20 for some damper-type configurations);   respectively  φh rj and φv rj are, respectively, the horizontal and vertical relative displacements between two ends of the jth damper associated with the fundamental   mode; φh i is the horizontal displacement of the fundamental mode at the ith floor. Note that the vertical mode displacements are known from the modal analysis carried out in Step 3 (Section 5.2.1.3).

5.2 Passive Damping Systems

Table 5.20 Vertical Magnification Factors for Various Damper Installation Schemes Magnification Factor of Vertical Direction ðGav Þi   sin θj

Damper Installation Type Diagonal

ðhj =bj Þ     cos θj2 cos θj4 2 θj1   cos θj1 1 θj2       cos θj2 cos θj1 1 θj3   2 sin θj3 cos θj1 1 θj2

Chevrona Up-Toggleb Bottom-Togglec

a

h is the interstory height and b is the length of a single bay. For this configuration, θ4 is the angle between damper direction and vertical direction. c For this configuration, θ3 is the angle between damper direction and vertical direction. b

In case the viscous dampers are nonlinear, the magnification factors, from Table 5.20, can be used with the following equation: Ci 5

•

P  2 ð2πÞ32α D12α 1D i m i φh i h     i11α ζ d P 22α T1 ψ ð G Þ φ 2 ð G Þ ah av h rj j j φh rj j

(5.81)

Here, D1D is the roof design displacement corresponding to fundamental mode (as calculated in Step 3 (Section 5.2.1.3)) and ψ is a constant defined in Chapter 4 (Section 4.1.1.1.1). Stiffness-proportional damping distribution. For linear viscous dampers, where a stiffness-proportional damping distribution is selected, the total added damping coefficient, Ctot , can calculated as follows (ASCE, 2017a): Ctot 5

T1

PN

2 i51 mi φ1i  2 2 j51 Kj ðGah Þj φ1j 2φ1;j21

4πKtot

PNd

ζd

(5.82)

where Kj is the jth lateral story stiffness (Step 2); Ktot is the sum of all the story stiffnesses; φ1j and φ1;j21 are, respectively, the horizontal modal displacements of the jth damper between the ith and ði 2 1Þth floor in the fundamental mode; and mi is the mass at the ith floor level. Consequently, the damping coefficient Ci for the ith story (this is the total damping needed per floor) can be directly calculated as: Ci 5 Ctot

Ki Ktot

(5.83)

If the dampers are nonlinear the following expression can be applied: Ci 5 Ctot

12α Ki  0:8u_ di;max Ktot

(5.84)

305

306

CHAPTER 5 Design procedures for tall buildings

•

Inversely linear distribution. If dampers are distributed as inversely linear along the height (see Fig. 5.13), the following expression can be used: Ci 5

T1

PN

2 i51 mi φ1i  2 2 j51 ηj ðGah Þj φ1j 2φ1;j21

PNd

4π



0:8u_di;max

12α

ζ d ηi

(5.85)

where Ci 5 ηi C denotes the total damping coefficient assigned to the ith story and ηi is known from Step 6 (Section 5.2.1.6). Note that damping coefficient evaluation can be accompanied or verified with the average values from a series of experimental tests (e.g., prototype and production tests (NEHRP, 2015)) during the design or construction phase (ASCE, 2017a; NEHRP, 2009, 2015) (see Chapter 7 for further details about these testing procedures). Damper power. Controlling the damper power is one of the most critical aspects in the design of viscous dampers. For example, for wind resistance the power needs to be controlled to prevent the damper from damage under lingering-high temperature. Considering an SDOF stimulated by a sine function (u 5 u0 sinðωtÞ) load, the energy dissipated of nonlinear damper, based on the damper force (Chapter 4 (Eq. 4.6))], is: ð ED 5

Fd dud 5

ð 2π=ω 0

cju_ d j11α dt 5 ψcωα u11α 0

(5.86)

where ψ is a constant estimated as shown in Eq. (4.20); ω is the angular velocity; and u0 is the amplitude of damper (often be set equal to 0.2
0.3 times maximum displacement of damper). This expression is equivalent to the one for linear dampers as shown in Eq. (4.9). Then the damper power is equal to: PD 5 ED f1

(5.87)

where PD is damper power in watts and f1 in general is the primary frequency of structure for the direction which has arranged dampers. Eq. (5.86) shows how the damper power is greatly influenced by the damping exponent. Hence, for damper used in wind resistance, the basic principle is the damping exponent cannot be too small. Moreover, for most dampers, their permitted power is less than 1 HP (different tonnage dampers have different permitted powers). Therefore, for α 5 0:3 the damper power will be overlarge if the damper is used for wind resistance, and this will be worse for smaller damping exponent coefficients. Balancing this by repeated iteration is one of the main contents of damper parametric design of high-rise buildings.

5.2.1.7.2 Viscoelastic dampers’ property estimation The stiffness and damping coefficients of viscoelastic dampers may depend upon the frequency and temperature of the dampers which are consistent with

5.2 Passive Damping Systems

the fundamental frequency of the building and the working temperature range (ASCE, 2017a). The stiffness and loss factor (damping) of each viscoelastic damper must be chosen based on the available viscoelastic material and the geometry of the damper (Chapter 4 (Section 4.1.1.1.2)). In general, this can be a trial-and-error approach (Christopoulos and Filiatrault, 2006) and it can be accompanied or verified with the use of a series of experimental tests (see Chapter 7) (ASCE, 2017a). As an assumption, the ambient temperature of the viscoelastic material can be assumed invariant, for example, 21 C, during earthquake events and the shear strain to remain constant at 100% (Marko et al., 2006). Moreover, it has been shown that a large decrease in the material stiffness occurs in the 0%
50% strain range, whereas in the 50%
200% strain range the stiffness remains approximately constant (Aiken et al., 1990). In general, depending on the distribution type of devices along the height (i.e., uniform-proportional and stiffness-proportional), two procedures are presented to preliminarily determine the viscoelastic damper properties (note that both procedures do not take into account the stiffness of the brace supporting the viscoelastic damper): •

Uniform damping distribution. In case the viscoelastic dampers are distributed uniformly along the total height or a part of total height, based on the supplemental viscous damping, ζ d , created by Nd discrete devices (Constantinou et al., 1998; Christopuolos and Montgomery, 2013), the damping coefficient assigned to ith story is: Ci 5

T1

PN

2 i51 mi φ1i  2 2 j51 ðGah Þj φ1j 2φ1;j21

4π

PNd

ζd

(5.88)

where T 1 is the fundamental period of the viscoelastic-damped building. To quickly estimate this period, the following expression is proposed by Ramirez et al. (2001): ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s 2ζ 12 d T 1 5 T1 η

(5.89)

where η is the loss factor of the viscoelastic material, as described in Chapter 4 (Section 4.1.1.1.2). For simplicity, ηD1 is suggested as an initial trial (Ramirez et al., 2001). Alternatively, this parameter can be more accurately calculated as shown in Appendix A (from Eq. 4.39) using the fundamental frequency, ω1 , and the operating design temperature of the building. A comprehensive database of viscoelastic material properties for a range of temperatures, strains, and frequencies is presented by Zimmer (1999). It is important to emphasize that the addition of viscoelastic dampers should not significantly alter the natural period of the building (Christopuolos and Montgomery, 2013). Consequently, based on the existing relation (Appendix A (Eq. 4.40)) between the stiffness and damping in viscoelastic

307

308

CHAPTER 5 Design procedures for tall buildings

materials, the viscoelastic damper stiffness property, kdi , of dampers associated with ith story can be estimated by: Ci kdi 5 η

•

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ω21 1 2 2ζ d =η

(5.90)

Stiffness-proportional damping distribution. The stiffness of viscoelastic dampers, kdi , added to the istory, can be set to be proportional to the story stiffness of the bare structure, Ki. The damper stiffness is then expressed, accounting for geometrical magnification factor ðGah Þi , expressed in Table 5.19, based on additional damping ratio, ζ d , at the fundamental mode of the structure, as (Soong and Dargush, 1997; Christopoulos and Filiatrault, 2006): 2ζ d Ki  kdi 5  η 2 2ζ d ðGah Þ2i

(5.91)

The viscoelastic damping coefficient can be calculated using Eq. (4.40) (as shown in Appendix A), and is also shown here: ηkdi Ci 5 qffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2

ω1 1 2 2ζ d =η

(5.92)

It should be noted that the response of the building may be assessed with multiple analyses, using limiting upper and lower bound amounts, for example, due to variation in temperature, for the stiffness and damping coefficients (ASCE, 2013). Designers can refer to Step 8 (Section 5.2.1.8.3) for more details about how to model changes in damper properties.

5.2.1.7.3 Displacement-dependent dampers’ properties estimation For displacement-dependent dampers the force is mainly a function of the relative displacement between two ends of the damper. Indeed, the response is independent of the relative velocity and the excitation frequency (see Chapter 4 (Section 4.1.1.3) for more details about this category of devices) (ASCE, 2017a; NEHRP, 2015). In this step, two kinds of displacement-dependent dampers are reviewed: •

Friction (coulomb) dampers are presented with three different damping distribution types (i.e., uniform-proportional, story shear-proportional, and inversely linear-proportional). Uniform-proportional to the story strength (shear) and triangle (inversely linear) distribution patterns of slips forces, in general, lead to a significantly better seismic performance (i.e., reduction in interstory drift and roof displacement). For example, Nadid et al. (2015) studied different pattern distributions for moment-resisting RC frames under seismic excitations retrofitted by friction dampers.

5.2 Passive Damping Systems

•

Hysteretic dampers with bilinear behavior (Liang et al., 2012) are reviewed. Two different distribution patterns are studied: uniform distribution along the height and story shear-proportional distribution. In addition, bracefriction (coulomb) damper systems (Lee et al., 2008a,b) are treated since they exhibit an elastic-perfectly plastic hysteretic model, similar to hysteretic dampers. Friction (coulomb) dampers. Depending on the distribution of friction dissipating devices along the building height, the frictional forces required for the design can be independently estimated using the effective supplemental damping ratio, ζ d . Assuming that the devices are pure friction dampers of a coulomb type (see Fig. 4.63 in Chapter 4), their mechanical behavior can be achieved ideally using zero velocity exponents, in the definition of the equivalent damping of nonlinear viscous dampers (as shown in the previous section). On this basis, different expressions are presented in the following, depending on damping distribution utilized. •

Uniform damping distribution. In case the damping force (coefficient) is associated with each of the Nd dampers, it is assumed to be identically distributed along the building height (Lee et al., 2008a,b), the following expression gives the preliminary design slip force of friction dampers: P d 2π3 ζ d D1M Ni51 mi φ21i Fdsi 5 2 PNm   T1 j51 fj φ1j 2 φ1;j21

•

(5.93)

where D1M is the maximum roof displacement under MCER associated with the fundamental mode that can be calculated from Eqs. (5.67) and (5.68). Story shear-proportional damping distribution. Friction dampers may be distributed proportional to the story shear forces, Vi (Pekcan et al., 1999a,b; Levy et al., 2001; Lee et al., 2008a,b). These forces can be estimated by applying a distributed (triangular) static lateral load on the structural system (Rao et al., 1995). Alternatively, if the story shear is unknown, the slip force with such a distribution type can be expressed by: Fdsi 5

P d 2π3 ζ d D1M Ni51 mi φ21i   Si PNm 2 T1 j51 Sj Gahj φ1j 2 φ1;j21

(5.94)

where the story shear force, Vi , assigned to ith story is proportional to the N P mi φ1i (Hwang et al., 2013). In alternative, the proportionalparameter Sj 5 i5j

ity of slip forces to story shears is defined based on the product of stiffness and peak interstory drift of the ith floor (Lee et al., 2008a,b). Additionally, the slip forces of friction dampers can also be distributed proportional to modal story shear forces (Levy et al., 2001).

309

310

CHAPTER 5 Design procedures for tall buildings

•

Inversely linear damping distribution. Nabid et al. (2015) proposed to distributed friction dampers based on an inverse-linear manner, where the slip force at the ith story, Fdsi , is proportional to the total slip force, Fds;tot , as: Fdsi 5 ηi Fds;tot P 2π3 ζ d1 D1M Ni51 mi φ21i Fdsi 5 2 PNm   ηi T1 j51 ηj Gahj φ1j 2 φ1;j21

(5.95) (5.96)

where ηi , called the proportionality factor, is assigned to ith story in such a way a linear distribution of slip forces can be obtained along the height. Hysteretic dampers. In general, the behavior of a hysteretic damper can be idealized with a bilinear damping mechanism (as reviewed in detail in Chapter 4, Section 4.1.1.2.1). Hence, when a damper is incorporated in a structure, the damping force versus relevant structural response can be modeled as a parallelogram (see Fig. 4.32 in Chapter 4). Hysteretic (bilinear) dampers can be characterized with the use of three independent parameters, that is, yielding displacement, Dy , yielding ratio (i.e., rα 5 k0p =k0 ), initial stiffness, k0 , and postyield stiffness, k0p . These parameters, shown in Fig. 4.32 (Chapter 4), should be specified from the list of manufacturer’s products (Liang et al., 2012). In particular, a parallelogram-type hysteretic model can also be obtained with a pure friction damper (Fig. 4.63 in Chapter 4) in series with an additional supporting system, connected to main structural members (see Fig. 5.16). In this case, the frictional force Fds in the damper, and the shear force, in brace Fb , are equal from a mathematical viewpoint. The hysteretic loop of a bracing-friction damper is then defined as an elastic-perfectly plastic model (Fig. 4.31 (Chapter 4)). Four steps are presented in the following; in order to preliminary design the hysteretic dampers/spring-friction damper systems for each story, given the supplemental damping ratio, ζ d , from Step 5 (Section 5.2.1.5) and fundamental Eigen properties (i.e., T1 and φ1 , Step 3 (Section 5.2.1.3)) of the building. •

Selection of yielding displacement: The yielding displacement, Dy , in hysteretic (bilinear) dampers plays a greater role in response reduction than the initial stiffness, k0 . The selection of the yielding displacement is frequently predetermined in accordance with the specific type of damping materials and in general, smaller values of Dy may give better results.

FIGURE 5.16 Spring-friction damper system.

5.2 Passive Damping Systems

•

•

Concerning spring-friction damper systems, yielding displacement (maximum preslip displacement) is related to the yielding properties of the brace supporting the friction damper. This can be chosen by the designer (Levy et al., 2001) or with reference to the manufacturer. Selection of yielding ratio: The value of this parameter is usually much less than unity. Designers may select the yielding ratio, rα , from the information provided by the manufacturer. When the yielding stiffness of devices is significantly small (i.e., k0 -0) this leads to an elastic-perfectly plastic hysteretic behavior which results in negligible yielding ratio (i.e., rα -0). This assumption is also valid for spring-friction damper systems. Determination of characteristic strength (f0 ): For simplicity, since the strain energy of the main structure is much larger than that stored in the devices, the latter one can be neglected once devices are yielded (Liang et al., 2012). Accordingly, assuming all the dampers are activated simultaneously, the following relation gives the approximate characteristic strength as: f0 5

•

P 2π3 ζ d D21M Ni51 mi φ21i     P m T12 Nj51 Gahj D1M φ1j 2 φ1;j21 2 Dy

(5.97)

where the maximum roof displacement, D1M , under MCER can be estimated from Eqs. (5.67) and (5.68) or from the result of a RSA using commercial software. Gahj is the geometrical factor for hysteretic dampers as given in Table 5.19. It is better to limit the characteristic strength, f0 , to a certain range, especially when the damper yielding stiffness, kd , is small in comparison to the structural stiffness. Moreover, f0 should not be too large or too small for vibration reduction (Liang et al., 2012). For example, in tall buildings (i.e., large fundamental period) the choice of a small characteristic strength will reduce the displacement more effectively (Liang et al., 2012). Concerning spring-friction damper systems (Fig. 5.16), the slip force, Fds , of the friction (coulomb) damper can be simply replaced by the characteristic strength, f0 , of hysteretic bilinear dampers (Fig. 4.32 (Chapter 4)). Determination of yielding stiffness and unloading stiffness: Given the damper parameters, the value of initial stiffness, k0 , and yielding stiffness, k0p , can be, respectively, calculated as: k0 5

f0 ð1 2 rα ÞDy

k0p 5 rα k0

(5.98) (5.99)

If the yielding stiffness, k0 , is less than a certain threshold, the structural vibration cannot be reduced. Hence, one can increase the k0 when the threshold is reached. It should be noted that, although increasing k0 is beneficial, it is usually determined by manufacturers; hence, designers do not have a selection range for the yielding stiffness (Liang et al., 2012).

311

312

CHAPTER 5 Design procedures for tall buildings

For brace-friction damper systems, rα 5 0 and accordingly, the brace stiffness (kb 5 k0 ) can be calculated as: kb 5 k0 5

f0 Fds 5 Dy Dy

(5.100)

Consequently, given the length, Lbj , of the brace, its cross-sectional area reads (Levy et al., 2001): Aj 5

kb Lbj EB

(5.101)

where EB is the elastic modulus of the brace material. Note that the above steps are suitable for hysteretic dampers where they are identical in every story (i.e., uniform damping distribution). For other kinds of damping distribution, the expression of characteristic strength is slightly different, depending on the type of distribution. For example, dealing with the story shearproportional damping distribution replacing the characteristic strength (f0i ) assigned to the ith story with the slip force (Fdsi ) in Eq. (5.94), the following expression is obtained to calculate the damper forces: f0i 5

P 2π3 ζ d1 D21M Ni51 mi φ21i     Sj j51 Sj Gahj D1M φ1j 2 φ1;j21 2 Dy

PNm 2

T1

(5.102)

5.2.1.8 Step 8: Construct damped structural model and perform structural analyses After the mechanical properties of dampers are preliminarily determined, as shown in the previous step, the mathematical model of the building structure, initially constructed in Step 3 (Section 5.2.1.3), should be updated by including the damping systems. Then, the NLTHA of the mathematical model of the structure and damping systems constructed in this step can be carried out for one of the following reasons: • •

To verify the responses predicted with the response-spectrum procedure as shown in Step 4.1 (Section 5.2.1.4.1). To mainly analyze the damped building responses, if the nonlinear procedure is chosen for the design.

The following points should be noted when dealing with the explicit damper structural modeling (ASCE, 2017a; NEHRP, 2015) in commonly available software: •

Structural system
NLTHA should be accomplished at both the service and MCER levels.
Inherent damping shall not be greater than 3% (NEHRP, 2015) (see Step 5, Section 5.2.1.5, and Chapter 3 (Section 3.2) for more details about inherent damping estimation).

5.2 Passive Damping Systems




•

Inherent eccentricity, due to asymmetry in mass and stiffness, can be accounted for in the model in MCER analysis.
Accidental eccentricity, consisting the center-of-mass by an amount equal to 5% of the diaphragm dimension separately in two orthogonal directions of building, should be considered in the model.
P-delta effects should be considered during modeling.
Depending on the material behavior, the hysteretic behavior of the structural members should be properly assigned during modeling.
Per the preliminary analysis (Step 6, Section 5.2.1.6), every structural element can be modeled as linear if capacity/demand ratio is less than 1.5.
The structures that have significant horizontal structural irregularity (in plan) should be modeled using a 3D representation.
If the floor diaphragms are not considered as rigid ones, the model may include representation of the diaphragm’s stiffness characteristics and relevant additional degrees of freedom.
The SSI may be included through direct analysis (through finite element modeling of the SSI) or substructure approach (through representing the stiffness and damping of the soil
foundation interface). More details about this is outside the scope of this publication but interested readers should refer to NIST (2012). Damping system
Results under the MCER analysis should be used to design the dampers.
If damper properties change with time and/or temperature, such behavior should be explicitly accounted for in the model (further details in Section 5.2.1.8.3 and in Chapter 4 (Section 4.1.1) for general material damper behavior).
In the direction of interest, depending upon the suitable positions in the structural system, at least two dampers at each side of the center of rigidity shall be modeled along all stories of the structural system (NEHRP, 2015); the mechanical properties of all the dampers in each story can be assumed identical.
Displacement-dependent dampers (e.g., friction and hysteretic) should be modeled accounting for the hysteretic behavior of the devices consistent with test data (see Chapter 7 for more details about testing devices).
The stiffness (e.g., in viscoelastic dampers and if any in viscous dampers) and damping properties of the velocity-dependent damping devices can be explicitly assigned using the values specified in Step 7 (Section 5.2.1.7).

The major commercially available software capable of modeling dampers are the following: CSI (2016a,b,c), Opensees (2017), Abaqus (Dassualt, 2017), Midas (Harpaceas, 2017), Seismosoft (2017), and Ansys (2017). Other softwares are commercially available but they are more specific to analyze the mechanical behavior of the single device, such as Comsol (2017) and MSC Nastran (2016).

313

314

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.17 Viscous/viscoelastic dampers modeling: (A) Kelvin model and (B) Maxwell model.

In the following, basics of dampers modeling are given. Interested readers should refer to the relevant literature on the subject for more details.

5.2.1.8.1 Modeling viscous/viscoelastic dampers Viscous/viscoelastic dampers can be modeled using the following two models: • •

Kelvin model (Fig. 5.17A). Mostly utilized for dampers with storage stiffness and weak frequency dependence, and discussed in detail in Chapter 4. Maxwell model (Fig. 5.17B). Mostly utilized for dampers with strong frequency dependency (this model ignores temperature dependence). CSI (2016b) suggests using this model for both viscous and viscoelastic dampers. The stiffness, kd , should be set to as large as possible. In order to guarantee this, it is recommended that the relaxation time (cd =kd ) to be an order of magnitude less than the loading time step, Δt. For example, a value of kd 5 100cd can be utilized.

Having the dampers always in series with the linkage of the damper-brace assembly will act as a Maxwell model. For this reason, the damper effectiveness is reduced as a function of structural properties and loading frequency (FEMA, 2006b). Nonlinear two-node link element (NLLINK), found in SAP2000 (CSI, 2016b), ETABS (CSI, 2016a), or OpenSees (2016) programs, can be utilized for modeling this type of dampers. This can be expressed as a Maxwell element with a dashpot and a spring in parallel (Fig. 5.17B (Christopoulos and Montgomery, 2013; DIS, 2017)). Note that SAP2000 (CSI, 2016b) has some numerical problems in case nonlinear viscous dampers with velocity exponents less than 0.4 are utilized.

5.2.1.8.2 Modeling friction and hysteretic dampers The nonlinear behavior of yielding and friction dampers can be modeled with a bilinear elastoplastic spring with hardening (Fig. 4.32 (Chapter 4)) and a bilinear rigid plastic (perfectly plastic, Fig. 4.31 (Chapter 4)) spring, respectively. For this

5.2 Passive Damping Systems

FIGURE 5.18 Plastic
Wen model.

purpose, a nonlinear link element available in most software, called Plastic
Wen (Fig. 5.18), can be used. This element has six DOFs and, for each DOF, the independent uniaxial plasticity properties (Bagheri et al., 2015) can be specified (Fig. 5.18).

5.2.1.8.3 Effects of variation in damper properties The design of dampers can include effects of environment (thermal conditions, aging, fatigue, and creep), production problems, and manufacturer test tolerances (velocity effects and first cycle effects). Hence, the nominal design properties specified and verified in previous steps may be modified using the property modification factors λ (ASCE, 2017a; NEHRP, 2009, 2015). Multiplying nominal design properties of dampers by property variation factors λmin and λmax (Section 5.1.3.3.3) leads to upper-bound (Equation 5.6) and lower-bound (Equation 5.7) design properties. Note that NEHRP (2015) allows to use an average value of λ, between λmin and λmax , for all the damping devices within a given type and size in lieu of each damper. These factors may be applied to whatever parameters (e.g., damping coefficient or velocity exponent) of the damper. The design temperature considered for the damper may range from annual low and high temperatures. It is suggested that the damper manufacturer is consulted when determining these factors. For a better understanding, a practical example is presented (NEHRP, 2015), illustrating the property variation factors for a nonlinear viscous damper with nominal properties C 5 128 kips:s=in and α 5 0:38. Fig. 5.19 shows the nominal force
velocity plot (filled curve) and the data (red points) obtained from a prototype test including: (1) ten full cycles performed at various amplitudes and constant temperature (21.1 C) and (2) three reversible cycles under various velocities and temperatures 4.4 C, 21.1 C, and 37.8 F. The maximum and minimum property variation factors can be estimated, respectively, in Eqs. (5.4) and (5.5) that give λmax 5 1:2 and λmin 5 0:8. Subsequently, the maximum and minimum

315

CHAPTER 5 Design procedures for tall buildings

500 400

Nominal –10%

300

+10% –20%

200

+20% Test data

Force (kips)

316

100 0 –100 –200 –300 –400 –500 –20

–15

–10

–5

0

5

10

15

20

Velocity (in/s)

FIGURE 5.19 Force
velocity relationship for a nonlinear viscous damper (NEHRP, 2015).

damping coefficients that must be considered from Eqs. (5.6) and (5.7) are Cmax 5 128 3 1:2 5 153:6 kips:s=in and Cmin 5 128 3 0:8 5 102:4 kips:s=in. This variation can be seen in Fig. 5.19 where it is shown the force
velocity relationship associated with 220% and 120% tolerance. The analysis of the building structure with dampers may be conducted separately, using Cmax and Cmin where the velocity exponent is fixed to 0.38. In this case, the first analysis may lead to larger damper forces and the second one results in less energy dissipation and larger drifts.

5.2.1.8.4 Load combinations The effects on the dynamic modification system and its components due to gravity loads and seismic/wind forces can be combined as shown from Eqs. (5.37)
(5.40) (Section 5.2.1.1.4). In the seismic load combinations, the horizontal seismic force, EL ; in the damping system shall be determined for the ELF or response-spectrum procedures as the maximum between the following three different stages (ASCE, 2017a): •

Maximum displacement rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi XNm E L 5 Ω0 ðFmSFRS Þ2 6 FDSD m

(5.103)

5.2 Passive Damping Systems

•

Maximum velocity rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi XNm ðFmDSV Þ2 m

(5.104)

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi XNm ðFCmFD Ω0 FmSFRS 1FCmFV FmDSV Þ2 6 FDSD m

(5.105)

EL 5

•

Maximum acceleration EL 5

• •

•

•

where: Ω0 is the overstrength factor based on the lateral force
resisting system chosen as given in Table 5.16 (Step 2, Section 5.2.1.2). FmSFRS is the force in an element of the damping system equal to the design seismic force for the mth mode. The definition of design force in the modal form Fim expressed in Eq. (5.55) is equal to FmSFRS . FDSD is the force in an element of the damping system necessary to resist maximum design seismic forces of displacement-dependent dampers, determined at displacements up to the design story drift, ΔiD , of the structure (refer to Step 5, Section 5.2.1.5). Note that FmSFRS and FDSD may differ depending upon the considered damper level. FmDSV is the force in the damping system required to resist design seismic forces for velocity-dependent dampers. For viscous dampers reads:    α FmDSV 5 max Fd;m 5 cdj ðGah Þj ri;mD

(5.106)

For viscoelastic dampers, this modal force can be expressed by:     FmDSV 5 kdj ðGah Þj Δi;mD 1cdj ðGah Þj ri;mD

•

(5.107)

where ri;mD is the interstory velocity for the ith story and the mth mode, at the design displacement. FC mFD and FC mFV are the force coefficient determined as follows:
Fundamental mode. The coefficients FC 1FD and FC 1FV are specified from Tables 5.21 and 5.22 using the velocity exponent, α, of viscous dampers, and damping is equal to the total damping minus the hysteretic one (see Step 4, Section 5.2.1.4).
Higher modes. The coefficients FC mFD and FC mFV can be specified from Tables 5.21 and 5.22 using the velocity exponent α 5 1, and total damping is determined for the displacement ductility, μ, in Step 4 (Section 5.2.1.4).

The coefficients CmFD and CmFV shall be taken as 1 for viscoelastic dampers, unless proved by test data or analysis. For intermediate values of α and μ interpolation is allowed for the values given in Tables 5.21 and 5.22.

5.2.1.9 Step 9: Check response acceptability The results from the RHA under wind/earthquake design, using a model consisting of both the structural and the damping systems, must be applied to recheck all the

317

318

CHAPTER 5 Design procedures for tall buildings

Table 5.21 Force Coefficient, FCmFD (ASCE 7-16 (ASCE, 2017a)) μ,1 Total Damping

α # 0:25

α 5 0:5

α 5 0:75

α $ 1:0

CmFD 5 1a

, 0.05 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 .1.0

1.00 1.00 1.00 1.00 1.00 1.00 1.00 1.00 1.00 1.00 1.00

1.00 1.00 0.95 0.92 0.88 0.84 0.79 0.75 0.70 0.66 0.62

1.00 1.00 0.94 0.88 0.81 0.73 0.64 0.55 0.50 0.50 0.50

1.00 1.00 0.93 0.86 0.78 0.71 0.64 0.58 0.53 0.50 0.50

μ $1 μ $1 μ $ 1:1 μ $ 1:2 μ $ 1:3 μ $ 1:4 μ $ 1:6 μ $ 1:7 μ $ 1:9 μ $ 2:1 μ $ 2:2

CmFD shall be taken as one for μ greater than the values shown.

a

Table 5.22 Force Coefficient, FC mFV (ASCE 7-16 (ASCE, 2017a)) Total Damping

α # 0:25

α 5 0:25

α 5 0:75

α $ 1:0

, 0.05 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 .1.0

1.00 1.00 1.00 1.00 1.00 1.00 1.00 1.00 1.00 1.00 1.00

0.35 0.44 0.56 0.64 0.70 0.75 0.80 0.83 0.90 1.00 1.00

0.20 0.31 0.46 0.58 0.69 0.77 0.84 0.90 0.94 1.00 1.00

0.10 0.20 0.37 0.51 0.62 0.71 0.77 0.81 0.90 1.00 1.00

elements of the structure. Any unacceptable response may include a dynamic instability, a nonconvergent analysis, a response that exceeds the allowable range of a deformation-controlled member, or the average strength of a critical force-controlled component (NEHRP, 2015). It is important to understand that this procedure may be iterative. Therefore, it might require an update of the damper parameters (i.e., going back to Step 5 (Section 5.2.1.5) and properties in Step 7 (Section 5.2.1.7)) in order to satisfy the criterion of the structural system and dampers. In particular, based on the appropriate load combination, as shown in Step 1.4 (Section 5.2.1.1.4), the following section briefly reviews the criteria for the structural design that shall be taken into consideration.

5.2 Passive Damping Systems

Table 5.23 Allowable Story Drift, Δa (ASCE, 2017a) Risk Category Structure

I or II

III

IV

Structure 4 stories or less (no masonry) with nonstructural elements detailed to accommodate story drifts Masonry cantilever shear wall structures Other masonry shear wall structures All other structures

0.025

0.020

0.015

0.010 0.007 0.020

0.010 0.007 0.015

0.010 0.007 0.010

5.2.1.9.1 Structural system For the structure-including dampers, the strength requirements should be satisfied using the demands obtained from the analysis with DE (NEHRP, 2015). The seismic force-resisting system needs to be designed for a strength that is not less than 75% of the strength without damping system (as already discussed in Step 4.1 (Section 5.2.1.4.1)). Further, considerations about member design of the main lateral force
resisting system, as well as the gravity system, are outside the scope of this book. Interested readers should refer to the relevant code on the subject (e.g., concrete (ACI, 2014) and steel (AISC, 2016a)).

5.2.1.9.2 Drift criteria According to design standards (e.g., ASCE (2017a)), the design story drift, ΔiD , (i.e., the drift under design ground motions) for any story i obtained from the time-history analysis (or RSA) of the structural system supported by dampers should satisfy the following criterion (ASCE, 2017a): ΔiD #

R Δa Cd

(5.108)

where R and Cd are listed, for the present structural system, in Table 5.16 and the allowable story drift Δa can be chosen from Table 5.23 with regard to the type of structure and risk category determined based on the occupancy of tall buildings (ASCE, 2017a). Furthermore, it is stated (NEHRP, 2015) that the peak story drift under MCER ground motion, ΔiM , may be checked with the use of the following criterion and drift limits specified in Table 5.23:  R ΔiM # min 0:03; 1:5 Δa ; 1:9Δa Cd

(5.109)

5.2.1.9.3 Acceleration criteria In order to be within acceptable limits with regard to human response to rather ordinary motions of buildings, horizontal accelerations of building with a specific

319

320

CHAPTER 5 Design procedures for tall buildings

return period (e.g., 1 or 10 years) are applied to the evaluation of habitability (e.g., ISO 10137 (2007)). In Chapter 3 (Section 3.4.2), different code requirements and relative perception thresholds were discussed in detail and the building response should follow these requirements. The peak accelerations of the target floor, computed from the history dynamic analysis, should not exceed the basic evaluation curve shown in Chapter 3, Fig. 3.24 (and relative Table 3.3) for the respective occupancy and the relevant code. It is important to note that the torsion acceleration should be verified also. This refers to the equivalent translational acceleration defined as the multiplication of the distance ðrt Þ from the center of torsion to the objective point and the angular acceleration, aθ ðtÞ, of the torsional vibration, that is, rt 3 aθ ðtÞ.

5.2.1.9.4 Damper criteria The following criteria shall be satisfied by the damping system: •

•

•

•

•

The dampers and their connections should resist the forces, displacements, and velocities from the MCER ground motions remaining in their elastic state (NEHRP, 2015). The hardware of all dampers (e.g., the cylinder of a piston-type device) and the connections between the devices and the structure must remain elastic under MCER. The damper components should be assessed with respect to their nominal strength in order to resist the seismic forces obtained from the nonlinear history procedure using the load combinations for the strength design, expressed in Eqs. (5.37) and (5.38). In case fewer than four energy dissipation devices per direction (or two on each side of the center of rigidity) are used, the maximum calculated displacement and velocity, under MCER, shall be increased by 130% (NEHRP, 2015). Damping device elements that cannot take any inelastic deformation capacity (force-controlled elements (NEHRP, 2015)) shall be designed for an increase of 20% of the MCER actions.

The forces and deformations provided from this analysis shall be given to the manufactures that will design accordingly the mechanical part of the damper device being utilized. Elements of the dampers are allowed to exceed strength limits under design loads if it is indicated by test or analysis where: • •

Inelastic responses do not adversely affect damper function. The design forces (QE ) in damper elements, computed by Eqs. (5.103)
(5.105) using Ω0 5 1, do not exceed the strength required to satisfy the load combinations expressed in Eq. (5.37).

5.2 Passive Damping Systems

A design review of the damping systems and relevant testing programs. After completion of the design calculations, a design review is recommended with the aid of an independent team of experts and experienced people in seismic analysis and passive energy dissipating systems. This step may require the review of the following (ASCE, 2017a; NERHP, 2015): • • •

Seismic criteria of the site comprising, for example, the development of site spectra and ground motion histories The preliminary design of the structural system and damping system, as well as design parameters of dampers The final design and all the analyses of the structural system and damping system

5.2.1.10 Step 10: Quality control, maintenance, and inspection requirements After the design of the building and damping components a quality control plan shall be written by the design professional to verify the following items (NEHRP, 2015): • • • •

Manufacturing process Preinstallation tests (e.g., prototype and production tests) to validate the designed damping properties and force
velocity
displacement relationships. Commissioning and system tuning Inspection/maintenance procedures A more detailed discussion on these aspects is given in Chapter 7.

5.2.2 STEP-BY-STEP PROCEDURE FOR MASS DAMPERS In Chapter 4 (Section 4.1.2), the essentials of mass damping systems, TMD, TLD, and TLCD were introduced. The same general step-by-step procedures shown in Fig. 5.8 have been utilized for this device category. In the following sections, first, some design considerations are reviewed and then the proposed step-by-step procedure for the design of mass dampers is reviewed in detail.

5.2.2.1 General design considerations 5.2.2.1.1 Tuned mass dampers The effectiveness of a TMD is determined by its basic design parameters: mass ratio (the ratio of TMD mass to the generalized mass of the building in its target mode of vibration) and TMD mass displacement. Depending on the target performance and the space constraints, a mass ratio in the range of 0.5%
2.0% is generally specified. Mass ratios higher than 2% can be used for cases where predicted wind-induced responses (i.e., accelerations) are significantly high and larger reductions of accelerations are required to bring them down to

321

322

CHAPTER 5 Design procedures for tall buildings

acceptable levels for occupant comfort. However, mass ratio increase has its own limits, which are driven by the limited gains in the TMD efficiency and the increased demand for its envelope space. Typically, a 30%
40% reduction in wind-induced acceleration responses of tall buildings can be achieved through the implementation of TMDs. Reductions as high as 50% can be practically achieved by high-efficiency TMDs which can generate a total equivalent damping ratio of 5%. Determining the size of the TMD mass and the space required for the entire TMD system is the first step of the TMD implementation. A close coordination with the design team is required at this stage to also determine the optimal location of the TMD room. Typically the modal height associated with the tuning frequency is the optimal location, which does not correspond to the top of the building. The top is just convenient from an architectural and coordination perspective. The mass is then computed based on the predicted accelerations at the uppermost occupied floor of the building for service-level winds and the dynamic properties of the building provided by the structural engineer (i.e., frequencies, mode shapes, mass distribution along the height of the building, and inherent structural damping ratio), to achieve the desired serviceability performance can be determined. Once the TMD mass is established, its anticipated peak displacement for design wind events is evaluated. This information is used as a basis to determine the envelope space requirements and develop the general arrangement. In general, several options are available to optimize the cost of the TMD mass, like choosing between different ballasts weight, such as steel, lead, concrete, and depleted uranium. Once the TMD concept has been developed and its envelope space requirements are fully coordinated with the architectural and structural design of the building, the performance of the TMD system must be evaluated through detailed time-domain wind and seismic response analysis for a wide range of wind speeds and directions, as well as earthquake intensities. The results of these analyses are used as a basis to determine the wind and seismic loads required for strength design and fatigue checks of the TMD components. Also, the results of the detailed performance analysis are used as a basis to determine the TMD
structure interface loads required for the structural design of the TMD supporting structure and additional supplementary damping devices if present.

5.2.2.1.2 Tuned liquid dampers The first step of the TLD sizing is to establish a suitable tank location and dimensions, through coordination with the design team. In situations where it is not possible to achieve the desired TLD mass with a single tank, multiple tanks can be employed by either stacking the tanks or locating them elsewhere on the floor plate. In these cases, superior TLD system performance can often be achieved by having each tank tuned to a slightly different frequency. The dimensions of the TLD must be determined from the natural frequency of the structural mode(s) to be controlled. The dimensions must be selected such that the TLD will function

5.2 Passive Damping Systems

satisfactorily over the expected range of as-built frequencies and return period wind events. The TLD must therefore have adjustable tuning and damping to enable device optimization after the final as-built frequencies are determined. This optimization must consider the variation in structural accelerations that will occur at different structural frequencies. Nonlinear phenomena, including liquid “spring-hardening,” and the incorporation of obstructions for the creation of liquid damping may alter the frequency of the tanks, should be taken into consideration as well. TLDs can often accommodate unusual tank geometries and remain effective when a building’s floor plan is not conducive to rectangular tank geometries. As the design develops, the performance of TLD systems must be validated through either nonlinear simulations that represent the nonlinear effects or scalemodel testing. Ultimate wind and seismic loads must be determined at the ultimate mean recurrence interval as dictated by the applicable building code. Due to the nonlinear behavior of TLDs, it is ill advised to apply factors to scale servicelevel loads up to ultimate loads. Since the TLD loads are oscillatory, it is necessary that all components are designed to resist fatigue loads that are expected at an acceptable mean recurrence interval. TLD tanks are generally constructed out of concrete and must be able to resist the peak dynamic pressures predicted under ultimate loading. The structural engineer is typically responsible for designing the concrete tank to resist the loads specified by the TLD designer. Careful attention must be paid to waterproofing, which is generally done using tank linings and/or concrete admixtures. If permitted by the applicable code, TLD tanks may serve the dual function as water storage for fire suppression. Implementation of the TLD requires careful coordination with the mechanical, electrical, and plumbing engineer to ensure provisions are made for water conditioning to prevent freezing and microbial growth. Drainage and fill provisions and interior tank access (for periodic inspection and maintenance) must also be coordinated. TLD installation is typically conducted by the general contractor with input from other trades as necessary. As construction approaches the floor that will house the TLD, the TLD designer should conduct frequency measurements on the partially completed building. Through coordination with the structural engineer, the measured frequencies are used to forecast the final as-built frequencies. By completing these measurements prior to the tanks being poured, it is possible to modify the TLD design if the forecasted frequencies are outside of the TLD tuning range, or finalize the tank dimensions for bidirectional systems. After the building is fully complete, the TLD designer must conduct final tuning and commissioning, in which the water depths and damping devices are finalized.

5.2.2.1.3 Tuned liquid column damper The first step of the TLCD design is the early coordination with the design team to find a geometry that will suit the design requirements. As a rule of thumb, TLCD generally requires that the long side of the tank stretches completely across the floor plan dimension, in the direction of the motion to be controlled. If more

323

324

CHAPTER 5 Design procedures for tall buildings

than one building vibration mode is to be controlled, additional tanks may be used, often arranged perpendicular to each other if necessary. The vessel can be narrower along its width, because it does not play any part in the determination of TLCD natural resonant frequency, since it cannot be bituned. Furthermore, upturned column dimensions need to be defined based on the column of water that they contain to correctly tune the TLCD to the as-built building structural properties. Therefore, the designer must anticipate some of the uncertainty inherent in estimating as-built structural frequencies, and have adjustable elements and/or a construction schedule that allows geometry adjustment in time for installation/pouring of the tank. Energy dissipation in a TLCD is generally provided through turbulenceinducing devices (i.e., screens) that induce nonlinear behavior to this type of liquid damper. Careful attention must be given during the design stage to ensure that sufficient adjustability is designed into the dissipation mechanism. This will permit optimal behavior to be realized when matching the as-built conditions of the building. The damping should be tested with a scale model early in the design stage to be certain of successful implementation. With sufficient design preparation, this dissipation can also be tested at full scale during the commissioning of the TLCD within the completed structure. The uncertainty of the as-built building frequencies, and fine-tuning of internal dissipation, can lead to a wide range of TLCD behaviors in very strong wind or seismic events. These ultimately mean that recurrence level events will cause fluid motion that exceeds the responses expected during less severe winds associated with serviceability performance. It is critical to evaluate the worst-case scenario of water column movement and the loads and pressures exerted by the moving water on its containment walls. These loads are calculated by the TLCD designer and are then typically given to the structural engineer to design the tank walls (and columns and ceiling). If chosen with care, the waterproofing system should last for decades. A costsaving decision to buy an inferior waterproofing system has repeatedly been seen to cause early problems, necessitating complete removal and replacement in only a few years. However, if chosen well, the level of maintenance required thereafter is little more than ensuring that the water level is kept at the proper depth. Microbial growth within the dark tank interior is only minor, and the MEP (mechanical, electric, plumbing) engineer should specify any necessary treatment, along with measures to prevent freezing. Fill and drain provisions, along with the access for periodic interior inspection, round out the list of other necessary considerations.

5.2.2.1.4 Available procedures for mass dampers The available design procedures for mass damping systems are listed in Table 5.24. While there is a wide literature on the subject, it is worth noticing that no particular design recommendations, in international standard codes, are available.

5.2 Passive Damping Systems

Table 5.24 Mass Damper: Available Design Procedures Available Methods TMD

TLD

TLCD

Soong and Dargush (1997) Rana and Soong (1998) Sadek et al. (1997) Chang and Qu (1998) Chang (1999) Chen and Wu (2001)

Fujino et al. (1992) Sun et al. (1995)

Sadek et al. (1998) Chang and Qu (1998) Chang (1999) Wu et al. (2005) Taflanidis et al. (2005) Wu and Chang (2006) Farshidianfar and Oliazadeh (2009) Shum (2009) Lee et al. (2011)

Connor (2003) Rasouli and Yahyahi (2002) Li and Liu (2002) Bakre and Jangid (2004) Miranda (2005) Gerges and Vickery (2005) Christopoulos and Filiatrault (2006) Kim et al. (2008) Ueng et al. (2008) Hoang et al. (2008) Ok et al. (2009) Moon (2010) Lu and Chen (2011a, 2011b) Aly (2012) Farghaly and Ahmed (2012) Moutinho (2012) De Angelis et al. (2012) Anh and Nguyen (2013) Özsarıyıldız and Bozer (2014) Elias and Matsagar (2014) Tuan and Shang (2014) Marian and Giaralis (2014)

Chang and Qu (1998) Yu et al. (1999) Banerji et al. (2000) Olson and Reed (2001) Tait et al. (2004) Tait (2008) Halabian and Torki (2011) Lee et al. (2011) Love and Tait (2012) Min et al. (2014) Chang (2015) Ruiz et al. (2016)

Quaranta et al. (2011) Di Matteo et al. (2014) Min et al. (2014)

The above-listed approaches could be generally applicable for building structures under both wind and seismic excitations. Many of these methods propose simplified relations to determine the optimal properties of dampers (e.g., mass, stiffness, and damping for TMDs; tank dimensions for TLDs; and tube characteristics for TLCDs). These are frequently established based on harmonic forces,

325

326

CHAPTER 5 Design procedures for tall buildings

harmonic ground motions, and white-noise ground accelerations of wind/seismic type. When the structure is subjected to more realistic excitations (wind events and earthquake ground motions), more complicated approaches (e.g., optimization algorithms) are required to optimize the damper properties. These methods, which are not treated in this book, include genetic algorithm (GA) (Hadi and Arfiadi, 1998; Singh et al., 2002; Marano et al., 2010; Mohebbi and Joghataie, 2011; Huo et al., 2013; Herve´ Poh’sie´ et al., 2015), bionic algorithm (Steinbuch, 2011), particle swarm optimization (Leung and Zhang, 2009), harmony search method (Bekdas and Nigdeli, 2011, 2013), ant colony optimization (Farshidianfar and Soheili, 2013), evolutionary operation (Islam and Ahsan, 2012), and charged system search (Kaveh et al., 2015). The majority of the present references do not offer step-by-step, yet simple, procedures, for the design of mass damping systems. Therefore, the design procedure presented in this section relies on different methods found in the literature (e.g., Soong and Dargush (1997); McNamara et al. (1999); Banerji et al. (2000); Connor (2003); Tait (2008); Christopoulos and Filiatrault (2006); Min et al. (2014); Tuan and Shang (2014); Chang (2015)). In the proposed step-by-step procedure, the estimation of optimal properties of damping devices is addressed mainly under harmonic forces, base harmonic accelerations, and artificial excitations (e.g., white-noise time histories of type wind and earthquake types). Hence, the procedure can be more accurately checked in the final phase using explicit modeling and more realistic load conditions (e.g., real ground motion accelerations). For this aim, commercial software packages (e.g., SAP2000 (CSI, 2016a); ETABS (CSI, 2016b)) can be used to easily simulate TMDs. For the case of TLDs/TLCDs, due to the nonlinear behavior of the liquid, it is more complicated to build a model using such a software; thus, one simplified choice is to model an equivalent TMD in lieu of TLDs or TLCDs with equivalent properties. It is important to note that the proposed design approach is potentially iterative, since the response criteria may be not satisfied during the first round of the design.

5.2.2.2 Step 1: Building and site categorization This step is identical to the one conducted for distributed dampers as described in Section 5.2.1.1.

5.2.2.3 Step 2: Select force-resisting system The second step of this method is to choose the appropriate lateral load
resisting structural systems for the tall building under study. Refer Section 5.2.1.2 for an in-depth review of the possible lateral force
resisting system.

5.2.2.4 Step 3: Building fundamental properties and preliminary structural analyses After the determination of the main structural system, modal properties, including frequencies, fm , mode shapes, φm , generalized mass, generalized stiffness, and

5.2 Passive Damping Systems

modal participation factors, should be estimated. For this aim, a numerical (linear) modal analysis with the use of available commercial software could help designers to compute these parameters. Enough modal periods and mode shapes should be calculated to obtain at least 90% mass participation of the actual mass in each horizontal direction of the building, as recommended by ASCE 7-16 (ASCE, 2017a). It should be noted that the effective stiffness for RC structural elements should be taken into account in the model in order to more accurately calculate the modal parameters (more details are provided in Section 5.1.3). Subsequently, the designer should analyze and design the bare structural system based on code requirements (ASCE, 2017a). From this analysis, the designer understand if the bare structural system can resist the lateral loads without requiring excessive structural members sizes and detailing in order to satisfy strength, drift, and code acceleration requirements. Moreover, for seismic loading, inelastic behavior is expected and this would induce damage to the structure causing reliability and economical concern as already explained in Chapter 3. Based on these analyses, the designer would understand the benefit of a dynamic modification system to be added to the main structural system selected in Step 2 (Section 5.2.2.3). In the following steps, the required procedure for the design of structures with mass damping technologies is reviewed in detail.

5.2.2.5 Step 4: Select a suitable analyses procedure As already discussed in Section 5.2.2.1.4, there are no codes available that regulate the design of this category of dynamic modification devices. Therefore, the only permitted analysis procedure (by ASCE (2017a)), for this category of devices, is NLTHA. The basics of this procedure were already discussed in Section 5.2.1.4.2 for distributed dampers and the same recommendations are to mass damper, as well.

5.2.2.6 Step 5: Select total target damping In this step, the inherent damping of the main structural and additional damping due to the isolated damping systems should be selected. For the inherent damping, a small value (usually less than 3%) should be chosen for all the modes, as recommended by ASCE (2017a) and NEHRP (2015). The designer should refer to Section 3.2 for an extensive discussion on the inherent damping and its selection. The total damping, ζ T , required for design purposes can be selected with respect to the dominant (wind-based or seismic-based) design type. For the windbased design, a wind-tunnel test can be adopted to estimate the required damping ratio, as recommended by McNamara et al. (1997). For this aim, with regard to a desirable response reduction percentage (e.g., roof lateral acceleration), the structural responses generated using the tunnel test, for various levels of damping, can be utilized to choose an appropriate effective damping. If the seismic-based design is of interest, response spectra can be used appropriately. In this case, the method proposed by Soong and Dargush (1997) and Christopoulos and Filiatrault

327

328

CHAPTER 5 Design procedures for tall buildings

(2006) is available for determining the effective damping ratio in the first mode by satisfying the following criteria: 8   < Γ1 SD ω1 ; ζ T # uN ðmaxÞ or  : Γ1 SA ω1 ; ζ T # u€ N ðmaxÞ

(5.110)

where Γ1 is the fundamental mode modal participation factor; ω1 is the fundamental mode natural frequency; ζ T , is the total effective damping associated with the fundamental mode; SD and SA are the design displacement and acceleration spectra, respectively; and uN ðmaxÞ and u€ N ðmaxÞ are the target (allowable) peak relative displacement and peak absolute acceleration at the roof level of the building, respectively. These thresholds are those which will be used in the final step to check the acceptability of the more accurate model results. More discussion on the selection of the total damping based on code requirements was given in Step 5 of the distributed damper procedures (Section 5.2.1.5).

5.2.2.7 Step 6: Damper type, configuration, and distribution 5.2.2.7.1 Step 6.1: Selection of the damper type In this section, the designer should decide which kind of mass damping systems (TMD, TLD, and TLCD) is most appropriate for the building design at hand. Several considerations need to be taken into account and some of them are the following: 1. Total cost of dampers. Several factors are contributed to the total cost of mass damper systems. In general, TMDs are slightly more expensive than liquid dampers because of their large mass (Wang et al., 2016). Moreover, the TLDs/TLCDs water tank can be useful for other emergencies, for example, fire-extinguishing water storage. Alternatively, the combination of tuned mass and liquid dampers is also proposed (e.g., Xu et al., 1992; Jae-Sung Heo et al., 2009; Wang et al., 2016) to utilize the economic advantages of TLCDs and the effectiveness of TMDs. 2. Directional control of response. It is easier to control the vibrations and responses in two directions, simultaneously, using TLDs (Sadek et al., 1998). 3. Required space to place dampers. Sloshing dampers usually requires less space than TMD because they do not require space for accommodating the stroke length (Sadek et al., 1998). 4. Type of dominant excitation force. The efficiency of mass damping systems in mitigating wind-induced response is extensively verified (see Chapter 4 for a detailed discussion). Matta (2013) reported that the effectiveness of TMDs decreases as the input duration shortens (e.g., pulse-like ground motions). An example of this could be for buildings located at near-field sites in the presence of ground motions with forward directivity or fling-step effects. Instead, sloshing liquid dampers, if appropriately designed, can be very

5.2 Passive Damping Systems

effective in controlling overall force, floor acceleration, and deformation responses of multistory building structures for broadband earthquake-type base excitations (Samanta and Banerji, 2012). 5. Modeling simplicity. It is well known that with the use of TMDs, the designer deals with physical properties (mass, stiffness, and damping) and it is an easy task to model these parameters explicitly. On the contrary, modeling of both TLDs and TLCDs is more complicated due to the nonlinear behavior of water (see Chapter 4 and Appendix A for a detailed discussion).

5.2.2.7.2 Step 6.2: Damper distribution Mass damping systems can be categorized with regard to the number of devices utilized in a building, such as single-tuned dampers (conventional devices) and multiple-tuned dampers (e.g., Iwanami and Seto (1984); Yamaguchi and Harnpornchai (1993); Fujino and Sun (1993); Abe and Fujino (1994); Kareem and Kline (1995); Sadek et al. (1998); Chang et al. (1998); Yalla and Kareem (2000); Li (2000), (2002); Min et al. (2005); Kim et al. (2006); Ashasi-Sorkhabi et al. (2014)). Single (conventional) mass damper. One of the advantages of conventional TMDs is related to its established technology as well as successful implementations in several tall buildings: for example, Taipei 101 (Section 8.2.3) (Soto and Adeli, 2013) (see Chapter 8 for several case studies). However, conventional TMDs can be tuned only for a certain structural frequency and this may be subject to uncertainties and to variations during strong motions. Furthermore, higher mode response control becomes a reliability problem since single TMD can control only the fundamental mode of vibration (Moon, 2010). Another main concern is the relatively high maintenance costs of such devices as they require special floor and various mechanical elements like springs, viscous dampers, and activation mechanisms. Similar considerations can be drawn for liquid mass dampers (see Chapter 4 and Appendix A for more details). Multiple mass dampers. Multiple-tuned mass dampers (MTMDs) can have two different distributions: horizontal distribution in one floor (usually the building top) (Elias and Matsagar, 2014) and vertical distribution along the height (Moon, 2010). The designer may eventually choose which kind of TMD configuration is appropriate based on the following considerations: •

Horizontal distribution. Superior reliability may be achieved by distributing MTMDs horizontally. Indeed, the system could work even if some of the devices are out of service or there are tuning errors (Moon, 2010). Kareem and Kline (1995) concluded that the MTMDs were most effective in controlling the motion of the primary system under random loading. In addition, they reported that the individual damping devices of MTMDs need less space than a massive single TMD, improving their constructability and maintenance. The essential characteristics of horizontally distributed MTMDs were investigated by some researchers (e.g., Yamaguchi and Harnpornchai

329

330

CHAPTER 5 Design procedures for tall buildings

•

(1993); Abe and Fujino (1994)), emphasizing their effectiveness and robustness for harmonically forced structural oscillations in comparison with an optimum single TMD. Vertical distribution. The designer usually chooses vertical distributed TMD since their efficiency is higher than single TMDs (Iwanami and Seto, 1984; Bergman et al., 1989; Xu and Igusa, 1992; Abe and Fujino, 1994; Moon, 2010). Jangid (1999) investigated the optimum parameters of MTMDs for an undamped main system. Results showed that the optimum damping ratio of the MTMDs decreases by increasing the number of TMDs and that the damping increases with the increase in the mass ratio. Moreover, the optimum bandwidth of an MTMD system increases with the increase in both mass and number of MTMDs. The most important studies carried out on this regards are the following:
Chen and Wu (2001) reported that the application of MTMDs installed at lower floors is more effective than upper floors for mitigating acceleration responses under seismic forces and that a single TMD is less effective. For displacement suppression, instead, the MTMDs almost behave identically to single TMDs.
Li (2002) emphasized the superior performance and robustness of optimum MTMDs against single TMDs.
Bakre and Jangid (2004) investigated the optimum parameters of MTMDs for suppressing the dynamic response of a base-excited damped main system using a numerical searching technique.
Lin et al. (2010) emphasized that the optimum MTMDs were not only effective in controlling the building responses but also successful in suppressing its stroke.
Moon (2010) concluded that using vertical distribution of MTMDs based on mode shapes can be advantageous in controlling not only the firstmode response but also those of higher modes. Moreover, this has a high potential of practical applications for tall building motion control. The zones to be considered with the vertically distributed TMDs for each mode of interest of the building can be specified based upon its mode shape (see Fig. 5.20 for the vertical distribution of TMDs for the first two modes (with the second one nonlinear) of a tall building (Moon, 2010)).

Regarding liquid dampers (TLDs and TLCDs), similar configurations to TMD can be utilized and, in literature, several studies have been conducted trying to look at the efficiency of multiple liquid dampers. A brief summary of the most important studies is the following: •

Sadek et al. (1998) reported that, although multi-TLDs (MTLDs) are not necessarily superior to single TLCDs, however, they are more robust regarding errors in the approximation of the structural parameters (Fujino and Sun, 1993; Gao et al., 1999).

5.2 Passive Damping Systems

FIGURE 5.20 Example of vertically distributed MTMDs based on mode shapes in a 60-story building. Adapted from Moon, K.S., 2010. Vertically distributed multiple tuned mass dampers in tall buildings: performance analysis and preliminary design. Struct. Design Tall Spec. Build. 19 (3), 347
366.

•

•

•

•

Gao et al. (2005) showed that a greater number of TLCDs in multi-TLCD systems enhances the efficiency of the system, but no significant enhancement is observed when the number of TLCD becomes more than five. Mint et al. (2004) demonstrated that the efficiency of MTLCDs is almost similar to that of the single TLCDs when there is no uncertainty in stiffness. MTLCDs show superior performance once stiffness uncertainty exists. Chang (2015) showed that if the required response reduction is less than 60%, then the single TLDs may be appropriate; in any other case MTLDs seem to be more robust. Mohebbi et al. (2015) proposed to utilize MTLCDs with different dynamic characteristics that can be located all in one floor or vertically distributed similarly to MTMDs (as discussed earlier).

5.2.2.8 Step 7: Damping system preliminary design In this step, the major parameters utilized for the design of mass damping systems are reviewed in detail. First the mass ratio between the damper and the structure is computed and then various optimal parameters per each device type are discussed in detail.

331

332

CHAPTER 5 Design procedures for tall buildings

5.2.2.8.1 Step 7.1: Mass ratio One of the most important design parameters for mass damping systems is the mass ratio, μ, which is defined as the ratio between the damper mass and the generalized mass of the main structure for the mode being suppressed. Depending on the type of mass damping systems (TMDs, TLDs, and TLCDs), different procedures can be utilized as explained in the following sections. TMD mass ratio. The general formulas were already reviewed in Chapter 4 for single (Eq. 4.23) and multiple (Eq. 4.125) TMDs. The main factor that influences the selection of the mass ratio is the required supplemental damping ratio. Indeed, using larger μ (i.e., large TMD mass) leads to a higher damping ratio (Bekdas and Nigdeli, 2013). However, this causes an increase in building weight; thus, the seismic forces could be amplified and excessive axial forces could occur in columns. Luft (1979) proposed the following simplified formulation based on the total and inherent damping, which was updated by Soong and Dargush (1997) and Chritopulos and Filiatrault (2006):  2 μ 5 16 ζ T 20:8ζ i

(5.111)

where ζ i and ζ T are the inherent and total damping, respectively, selected in Step 5 (Section 5.2.2.6). Further insights in the selection of the mass ratio can be found in the literature. Marano and Chiaia (2010) reported that TMD with a small mass ratio controls the response via resonance, leading to a rather large movement in relation to the main structure. The control mechanism differs for large TMD mass ratio. Indeed, Bekdas and Nigdeli (2013) recommended to have a minimum mass ratio for high-rise buildings. Moreover, the designer should take into consideration physical limitations, for example, available space at the floor of interest in preselecting the mass ratio. McNamara et al. (1997) stated that the mass ratio can be selected from the available charts of mass ratio versus the equivalent damping ratio, ζ T . According to this reference, μ approximately equals 2% for the first mode. Besides, Connor (2003) reported that for most applications, the mass ratio should be less than 5%. Therefore, based on these studies, as a general rule of thumb, the mass of a TMD can be selected around 0.25%
1.0% of the fundamental modal mass of the building (Chey, 2007). TLD mass ratio. The TMD mass ratio is defined as the ratio between the damper and the structural mass (Chapter 4 (Section 4.1.2.2)) (Tait et al., 2004a,b; Tait, 2008). Banerji et al. (2000) have shown that if the mass ratio is less than 1% and the inherent damping of the main system is greater than 2%, then the TLD is not going to be effective as a control device. TLCD mass ratio. The mass ratio for TLCDs can be computed as shown in Chapter 4 in Eq. (4.28) and (4.29) for single and multiple TLCDs, respectively. Sadek et al. (1998) state that the mass ratio, μ, should be selected based on the tradeoff between the response reduction, cost, space, and weight of the TMDs. Based on these aspects they recommended that the practical range of μ is from 0.5% to 4%. Similarly, Shum (2009) recommended a range between 0.5% and 5%.

5.2 Passive Damping Systems

5.2.2.8.2 Step 7.2: Determination of TMD design parameters Optimal tuning frequency and damping ratio. Generally speaking, the design of TMDs involves the selection of three parameters: mass, damping, and stiffness. The ideal values of these parameters depend on the optimal properties of TMDs: mass ratio, μ, frequency ratio, f d;opt , and damping ratio, ζ d;opt . These properties may be identified with respect to four criteria: the damping existence in the main structure; the type of excitation force (i.e., harmonic or random); the location of the excitation (i.e., directly on the primary structure or on at its base); and the type of control criteria (e.g., displacement or acceleration). Simplified relationships can be utilized based on two categories: (1) undamped and (2) damped systems. 1. Simplified relations for undamped systems with single TMD. In the existing literature, there are available several relationships for undamped structures (i.e., negligible inherent damping). Table A.1 (Appendix A) lists the available relations with regard to the type and position of the excitation, and also on the control criteria to optimize. 2. Simplified relations for damped systems with single TMD. Similar to the simplified relations defined before for undamped systems in the case of damped structures, Table A.2 (Appendix A) shows a list of available relations with regard to the type and position of the excitation, and also on the control criteria to optimize. In alternative to the relations presented in Table A.2 (Appendix A), design charts are available (e.g., Connor (2003) numerically developed the curves (see Figs. 4.A12 and 4.A.13 (Appendix A) (Connor, 2003)). Determination of mass, stiffness, and damping coefficients of TMDs. With regard to the choice of single or multiple TMDs for controlling the building response, the mechanical properties of TMD systems can be correspondingly determined. •

•

Single (conventional) TMDs. After determining the optimal tuning frequency, f d;opt , and damping ratio, ζ d;opt , of the single TMD, given the mass ratio from Step 7.1 (Section 5.2.2.8.1), the physical mass, md , stiffness, kd , and damping constant, cd , of the TMD can be calculated as follows: md 5 μmm

(5.112)

2 kd 5 f d;opt ω2m

(5.113)

cd 5 2ζ d;opt f d;opt ωm

(5.114)

where m denotes the mode of vibration of the primary system for which the TMD is tuned. In this stage, the mechanical parameters of the TMD determined may be checked by the device manufacturer. MTMDs. Manufacturing of MTMDs with uniform stiffness is simpler than those with varying stiffnesses (Xu and Igusa, 1992; Elias and Matsagar, 2014).

333

334

CHAPTER 5 Design procedures for tall buildings

Based on this assumption, the stiffness of TMDs can defined as (Bakre and Jangid, 2004): kd;1 5kd;2 5    kd;Nd 5kd

(5.115)

Therefore, according to the assumptions and the formulations proposed by Bakre and Jangid (2004), the mass of jth TMD can be expressed by: md;j 5

md;1 ω21 ωd;j 2

(5.116)

where md;1 5

μm 11

ω21 ω22

1  1

ω21 ωNm 2

(5.117)

and    Nd 1 1 ω d;opt ; ωd;j 5 f d;opt ωm 1 1 j 2 2 Nd 2 1

j 5 1; :::; Nd

(5.118)

where f d;opt and ω d;opt are presented in Table A.3 (Appendix A). Thus the identical stiffness for every TMD is given by: kd;j 5md;j ω2d;j 5kd

(5.119)

Instead, the optimal damping coefficient of jth TMD can then be determined as: cd;j 5 2md;j ζ d;opt ωd;j

(5.120)

where ζ d;opt is the damping ratio that is kept constant for all the MTMDs (see Table A.3 (Appendix A)). In case the stiffness and mass of each TMD are different, Elias and Matsgar (2014) proposed to use the following relationships: kd;j 5  1 ω2d;1

μ 1

PNm m51 1 ω2d;2

mm

;

1...1 md;j 5

j 5 1; . . .; m

(5.121)

1 ω2d;n

kd;j ω2d;j

(5.122)

where mm is the generalized modal mass of the building for the mth mode. The damping coefficient can then be estimated with Eq. (5.120).

5.2.2.8.3 Step 7.3: Determination of design parameters of TLD Optimal tuning frequency and damping ratio. To get the optimal performance, one can select the ωd equal to the dominant frequency of the main structure, ω, leading to the tuning ratio to be unity, that is, f d;opt 5 1 (Banerji et al., 2000). The

5.2 Passive Damping Systems

other approach to get the optimal tuning frequency, while the main system is undamped, is to use (Warburton, 1982a,b; Tait and Deng, 2010): pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 μ=2 f d;opt 5 11μ

(5.123)

In this case, the effective optimum damping ratio of the structure-TLD system can be estimated as follows (Tait, 2008; Tait and Deng, 2010): 1 ζ d;opt 5 4

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi μ 1 μ2 1 1 3μ=4

(5.124)

TLD tank dimensions. Based on the geometrical shape of TLD tank (e.g., rectangular, circular; see Chapter 4 for a deeper discussion on which geometry is more efficient), the various design dimensions can be calculated as follows:   • Liquid depth ratio rTL 5 dTL =LTL . This parameter denotes the ratio between the still water depth in the tank and the tank length. The practical value for the liquid depth ratio may be selected as close to the shallow water depth limit, that is, 0:15 (Banerji et al., 2000) (see Chapter 4, Section 4.1.2.2, for a detailed discussion on the water level in a TLD tank). The robustness of a TLD can be improved by selecting smaller liquid depth ratios (Tait, 2008; Tait and Deng, 2010). • Tank length ðLTL Þ. Given the liquid depth ratio, using the definition of linear water sloshing frequency (Banerji et al., 2000; Tait, 2008), the following expression gives the tank length in the direction of the sloshing motion: g

LTL 5

2

4πf d;opt fx2

tanhðπrTL Þ

(5.125)

where g is the gravitational acceleration and fx is the fundamental mode frequency of the structure in the x-direction. Note that the use of damping screens inside the tank guarantees the validity of this expression over a larger range of input excitation amplitudes (Tait, 2004). The other main direction length, bTL , can be determined similarly as follows: bTL 5

• •

g 2

4πf d;opt fy2

tanhðπrTL Þ

(5.126)

Note that bTL 5 LTL if the fundamental mode frequency of building is identical in two orthogonal directions. Water depth ðdTL Þ. After determining the tank length ðLTL Þ, the still water length can be easily calculated by dTL 5 rTL LTL . Number of tanks ðNd Þ. Having determined the nominal dimensions of the TLD tank, it is now possible to estimate the required number of tanks (Banerji et al., 2000): Nd 5

μm LTL  bTL  dTL  ρF

(5.127)

where m is the mass of primary structure and μ is the selected mass ratio.

335

336

CHAPTER 5 Design procedures for tall buildings

Equivalent mechanical properties of TLD. It may be possible to model the behavior of TLD systems with an equivalent TMD system; otherwise, complex fluid nonlinear analyses are required (see Section 4.1.2.2.2 (Chapter 4) and Appendix A.6). In the past decades, several simplified approaches have been proposed, as follows: •

•

•

•

•

•

Sun et al. (1995) proposed equivalent mass, stiffness, and damping of the TLD using an SDOF-TMD analogy and the experimental results are measured for rectangular, circular, and annular tanks, subject to harmonic base excitation. Yu et al. (1997) modeled TLD as an equivalent TMD with nonlinear stiffness and damping. The model described the behavior of TLD under a wide range of excitation amplitude. Empirical expressions were developed (Yu et al., 1999; Olson and Reed, 2001) for determining the properties of an equivalent nonlinear TMD model that captures the behavior of a TLD system under a variety of loading conditions. The properties of an equivalent amplitude-dependent TMD, with equal energy dissipation as a TLD equipped with damping screens, were evaluated experimentally using shake table tests (Tait et al., 2004a) and the performance of this semiempirical amplitude-dependent model was verified as well (Tait et al., 2004b). An equivalent mechanical linearized structure-TLD system with equivalent mass, stiffness, and damping was developed by Tait (2008) for both sinusoidal and random excitations; and was validated using experimental tests. Ruiz et al. (2016) proposed a TLD with a floating roof (TLD-FR), consisting of a traditional TLD with the addition of a floating roof. For this kind of isolated damper, a dynamic behavior based on linear TMD can be better captured.

In the following, several studies are reviewed for estimating the equivalent properties for rectangular and circular tanks. Rectangular tanks Tait (2008) provided the expressions for the equivalent mass and stiffness, corresponding to the fundamental sloshing mode of the TLD system with a rectangular tank as follows: 8ρF bTL LTL tanhðπrTL Þ π3

(5.128)

8ρF bTL LTL g tanh2 ðπrTL Þ π2

(5.129)

md;eq 5 kd;eq 5

The damping coefficient can be estimated for random and sinusoidal excitation as shown in the following (Tait, 2008): •

Random excitation for a given value of an equivalent response motion, σr , (Tait, 2008): cd;eq 5 Cl

16ρF bTL LTL π3

rffiffiffiffiffi 32 tanh3 ðπrTL Þδn Ξdm ωdm σr π3

(5.130)

5.2 Passive Damping Systems

100

Cd CI

Cd, Cl

10

Modified Cd

1

Modified CI 0.1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

S

FIGURE 5.21 Drag coefficient values for slat screens (Tait et al., 2005).

where Cl is the loss coefficient of the screens inside the tank; this parameter can be estimated from Fig. 5.21. In the figure, the solidity parameter, STL , is defined as (Tait, 2008): STL 5

ATL bTL dTL

(5.131)

where ATL is the area of the (submerged) screen normal to the flow. Note that for solidity ratios less than 0.3, the modified curves (dashed lines) shown in Fig. 5.21 should be used (Baines and Peterson, 1951). The frequency of the mth mode of the sloshing water is expressed by (Tait, 2008): ωd;m 5

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi mπg tanhðmπrTL Þ LTL

(5.132)

The parameter δn (damping due to the screen-induced losses related to the horizontal component of flow) in the mth mode of the sloshing water can be obtained as (Tait, 2008): δm 5

1 1 1 3 sinh2 ðmπrTL Þ3

(5.133)

The damping due to screen locations in the tank, Ξd;m , can be calculated for the mth mode of the sloshing water by (Tait, 2008): Ξd;m 5

Ns X j51

 mπxj 3 sin LTL

(5.134)

337

CHAPTER 5 Design procedures for tall buildings

where xj is the location of the jth screen from the left-hand side edge of the tank and Ns is the number of screens that should be selected by the designer. The parameter σr , root mean square (RMS) relative response motion of the equivalent mechanical linear model is given by (Tait, 2008): sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 11μ σr 5 σs 2μ 1 3μ2 =2

•

(5.135)

where σs is the target structural response level that can be estimated from Fig. 5.22 with the use of optimum effective damping ratio ζ d;opt expressed in Eq. (5.124). In the figure the value of σs;target is the structural response amplitude resulting in the highest level of effective damping for the structure-TLD system. The equivalent damping coefficient for the sinusoidal excitation is given by (Tait, 2008): cd;eq 5 Cl

256ρF bTL LTL tanh3 ðπrTL Þδm Ξd;m ωd;m xr 3π5

(5.136)

where xr is the displacement variable associated with the equivalent linearized mechanical model of the TLD system.

4.0

3.0

ζeff (%)

338

2.0

1.0 Analytical Experimental 0.0 0.0

0.5

1.0 σs/σs,target

1.5

2.0

FIGURE 5.22 Measured and predicted effective damping versus normalized target structural response level (Tait, 2008).

5.2 Passive Damping Systems

Alternatively, Chang and Qu (1998) proposed to calculate the equivalent mass, stiffness, and damping of the rectangular TLD as follows: mdeq 5 ρF bTL LTL dTL 8ρF bTL LTL g tanh2 ðπrTL Þ π2 rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi g tanhðπrTL Þ cdeq 5 2ζ d;opt ρF bTL LTL dTL π LTL kdeq 5

(5.137) (5.138) (5.139)

Circular tank For circular TLD systems, the following relationship can be utilized (Chang and Qu, 1998) for determining the main parameters: md;eq 5 ρF πL2TL dTL

(5.140)

kd;eq 5 0:419ρF gπL2TL tanh2 ð1:84rTL Þ

(5.141)

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi g tanhð1:84rTL Þ cd;eq 5 2πζ d;opt ρF L2TL dTL 1:84 LTL

(5.142)

5.2.2.8.4 Step 7.4: Determination of TLCD design parameters Optimum tuning frequency and damping ratio. The design parameters of single TLCDs, that is, tuning ratio and head loss coefficient, and of multiple TLCDs (MTLCDs), that is, central tuning ratio, tuning bandwidth, and number of damper groups, can be selected based on the following theory. •

Single TLCDs Tuning frequency (f d;opt ). Under white-noise base-acceleration ground motions, the optimum tuning frequency can be calculated as (Sadek et al., 1998): pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 f d;opt 5 11μ

(5.143)

Instead, in the case of earthquake excitation, the optimum tuning frequency can be estimated as follows (Chang, 1999): pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 μ 2 3ΨTLCD =2 f d;opt 5 11μ

(5.144)

 2 where ΨTLCD 5 μ bTLCD =LTLCD is called the efficiency index that is proportional to the ratio between the width, bTLCD , and total length, LTLCD , of the TLCD tube. Chang (1999) suggested the values of ΨTLCD 5 0:6μ, ΨTLCD 5 0:8μ, and ΨTLCD 5 1:0μ. Similarly, Sadek et al. (1998) recommended a value of ΨTLCD 5 0:8μ. Subsequently, the relative optimum damping ratio may be estimated as follows (Chang, 1999): 1 ζ d;opt 5 2

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 1 μ 2 5ΨTLCD =4   ð1 1 μÞ 1 1 μ 2 3ΨTLCD =2

(5.145)

339

340

CHAPTER 5 Design procedures for tall buildings

Note that the inherent damping of the main systems is neglected; if the main system is damped, the optimum relations proposed for TMDs in Table A.2 (Appendix A) can be used depending on the type of excitation and its position. For wind loading excitation, the optimum tuning frequency may be expressed by (Chang, 1999): f d;opt 5

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 μ 2 ΨTLCD =2 11μ

(5.146)

Based on the latter equation, Shum (2009) proposed a similar expression, while the multiplier of ΨTLCD is unity in lieu of 1/2. Subsequently, the relative optimum supplemental damping ratio can be expressed by: 1 ζ d;opt 5 2

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ΨTLCD 1 1 μ 2 ΨTLCD =4   ð1 1 μÞ 1 1 μ 2 ΨTLCD =2

(5.147)

Head loss coefficient ðηd;opt Þ. The other design parameter of TLCDs is the head loss coefficient which depends on the expected ground acceleration and the orifice opening ratio (area of opening to cross-sectional area of tube) (Sadek et al., 1998). Note that the head loss coefficient is used to select the orifice opening of the TLCD. For ηd 5 0; the orifice is fully open and for ηd 5 N, the orifice is closed. The value of ηd can be defined in terms of orifice opening ratio, as explained by Blevins (1984), or it can be determined from experiments for specific orifice shapes and sizes. Therefore, the head loss coefficient can be expressed by (Sadek et al., 1998): ηd;opt 5

3:58μ a=u€ g;max

(5.148)

The practical ranges of the mass and inherent damping ratio are: ratio is 0:05 # μ # 0:04 and 0:02 # ζ i # 0:05, respectively; u€ g;max denotes the maximum ground acceleration, as a multiplier of gravity acceleration, g. The range of this parameter may be assumed as 0:05g # u€ g;max # 1g. Wu et al. (2005, 2009) proposed design tables for TLCD parameters including the values of optimum design parameters (tuning frequency, head loss coefficient, and effective supplemental damping ratio). These values are extracted for both harmonic and white-noise excitations, based on various primary mass inherent damping ratio, ζ i , mass ratio, μ, and ratio of the horizontal length to total length of the liquid column of TLCD, rTLCD . Table 5.25 lists the optimal values for harmonic loading for inherent damping of the main mass 1%. Similarly, Table 5.26 gives the corresponding numbers for white-noise loading cases. Similarly, Shum (2009) presented design tables for designers in order to estimate the optimum values of tuning frequency and head loss coefficient.

5.2 Passive Damping Systems

Table 5.25 Harmonic Loading and Inherent Damping (Wu et al., 2009, 2009) ζ i 5 0:01 ηd;opt ð1022 Þ μ

r TLCD

f d;opt

95%a

100%b

95%c

ζ d;opt

0.01

0.5 0.6 0.7 0.8 0.9 0.5 0.6 0.7 0.8 0.9 0.5 0.6 0.7 0.8 0.9 0.5 0.6 0.7 0.8 0.9 0.5 0.6 0.7 0.8 0.9

0.9930 0.9923 0.9915 0.9906 0.9897 0.9866 0.9854 0.9839 0.9822 0.9804 0.9805 0.9786 0.9765 0.9741 0.9714 0.9744 0.9721 0.9693 0.9661 0.9626 0.9685 0.9656 0.9623 0.9584 0.9541

6.073 6.898 7.722 8.535 9.365 15.465 17.773 20.052 22.494 24.873 26.998 31.289 35.596 39.966 44.378 40.181 46.835 53.524 60.220 67.114 54.867 64.132 130.509 83.069 92.721

9.519 10.687 11.841 12.849 14.023 23.523 26.563 29.275 33.127 36.385 40.511 46.472 52.253 58.169 64.132 59.503 68.936 78.000 86.624 96.218 81.047 93.760 106.811 119.471 132.435

12.543 13.765 15.031 16.320 17.632 30.091 33.740 37.467 41.220 45.016 51.051 57.869 64.770 71.748 78.790 74.714 85.288 95.976 106.753 117.689 100.689 115.500 73.576 145.660 161.011

2.64 2.99 3.34 3.69 4.05 3.35 3.85 4.34 4.84 5.34 3.89 4.49 5.10 5.71 6.33 4.34 5.03 5.74 6.44 7.15 4.72 5.50 6.29 7.08 7.87

0.02

0.03

0.04

0.05

a

5% degradation (from the left) from the optimal head loss coefficient on each side. The case related to the optimal head loss coefficient. c 5% degradation (from the right) from the optimal head loss coefficient on each side. b

•

MTLCDs Optimum parameters of MTLCDs are conceptually similar to MTMDs, where the important design parameters are the frequency range of the devices and their damping ratio (Yamaguchi and Harnpornchai, 1993; Kareem and Kline, 1995). Three parameters can be proposed to characterize an MTLCD system: central tuning ratio, fd;0 , tuning bandwidth, Δfd , and the number of TLCD groups, Nd (Sadek et al., 1998). The central tuning ratio can be computed as follows: fd;0 5

fd;Nd 1 fd;1 2

(5.149)

341

342

CHAPTER 5 Design procedures for tall buildings

Table 5.26 White-Noise Loading and Inherent Damping (Wu et al., 2005, 2009) ζ i 5 0:01 ηd;opt ð1022 Þ μ

r TL

f d;opt

95%a

100%b

95%a

ζ d;opt

0.01

0.5 0.6 0.7 0.8 0.9 0.5 0.6 0.7 0.8 0.9 0.5 0.6 0.7 0.8 0.9 0.5 0.6 0.7 0.8 0.9 0.5 0.6 0.7 0.8 0.9

0.9942 0.9939 0.9935 0.9931 0.9926 0.9886 0.9880 0.9873 0.9865 0.9856 0.9831 0.9823 0.9812 0.9800 0.9788 0.9778 0.9766 0.9753 0.9738 0.9721 0.9725 0.9711 0.9695 0.9676 0.9655

1.702 1.853 1.987 2.110 2.225 3.983 4.323 4.632 4.917 5.183 6.528 7.083 7.589 8.055 8.490 9.261 10.048 10.765 11.426 12.040 12.141 13.172 14.110 14.975 15.776

3.474 3.623 3.771 3.918 4.063 7.545 7.955 8.351 8.731 9.096 11.979 12.697 13.383 14.030 14.648 16.677 17.735 18.734 19.675 20.564 21.586 23.006 24.337 25.586 26.760

7.445 7.369 7.398 7.482 7.597 14.769 15.033 15.396 15.805 16.234 22.557 23.246 24.017 24.814 25.614 30.701 31.872 33.097 34.325 35.526 39.134 40.827 42.541 44.224 45.851

2.04 2.29 2.53 2.78 3.03 2.55 2.90 3.25 3.60 3.95 2.94 3.38 3.80 4.24 3.475 3.28 3.77 4.27 4.78 5.29 3.57 4.13 4.69 5.26 5.81

0.02

0.03

0.04

0.05

a

The case related to the 5% degradation (vicinity) from the optimal head loss coefficient on each side. The case related to the optimal head loss coefficient.

b

where fd;1 and fd;N refer to the tuning frequency of the first and Nd th group of TLCDs, respectively. An important assumption  of thisdefinition is that the difference between adjacent tuning ratios, fd;j11 2 fd;j , are assumed to be constant. Sadek et al. (1998) reported that for the best efficiency in response reduction, the central tuning ratio, fd;0 , of an MTLCD system should be tuned to the (dominant) natural frequency of the main structure. The tuning bandwidth frequency is defined by: Δfd 5

fd;Nd 2 fd;1 fd;0

(5.150)

5.2 Passive Damping Systems

Sadek et al. (1998) recommended that Δfd should be 0.125, 0.1, 0.05, and 0.025 for mass ratios 0.04, 0.02, 0.01, and 0.005, respectively. Given the value of fd;0 and Δfd from Eqs. (5.149) and (5.150), respectively, the tuning ratios, fd;j , associated with jth group of TLCDs is calculated by: fd;j 5 fd;1 1 ði 2 1Þ

where

fd;Nd 2 f1 Nd 2 1

8 0 1 > > Δf > d > @ A > < fd;1 5 fd;0 1 2 2 > fd;0 > > f 5 ð2 1 Δfd Þ > > : d;Nd 2

(5.151)

(5.152)

Tube dimensions. In this section, the geometrical parameters (i.e., liquid column length, cross-sectional area of the tube, tube width to liquid length ratio, and tube width) of single TLCDs and of MTLCDs can be determined by the designer (Sadek et al., 1998). •

Single TLCDs Liquid column length ðLTLCD Þ. Given the optimum tuning frequency, f d;opt , from Eqs. (5.143), (5.144), or (5.146), the liquid column length, LTLCD , can be determined by the following (Sadek et al., 1998): LTLCD 5

2g 2 ω2 fd;opt

(5.153)

where ω is the natural frequency of the main mass. Cross-sectional area of the tube ðATLCD Þ. Given the mass ratio from Step 7.1 (Section 5.2.2.8.1) and the liquid length, LTLCD , the cross-sectional area of the tube of TLCDs can be easily obtained as: ATLCD 5

μm ρF LTLCD

(5.154)

  Tube width to liquid length ratio rTLCD 5 bTLCD =LTLCD . For this design parameter a range from 0.1 to 0.9 is suggested by Sadek et al. (1998) that the larger the ratio the larger the response reduction. Moreover, Sadek et al. (1998) suggest an optimal value of 0.80 for practical design. Alternatively, Shum (2009) proposed values between 0.7 and 0.9 for an economical design. However, it is important to remember that Sun et al. (1995) reported that increasing this parameter enlarges the RMS displacement of the structure. To get the best performance of TLCDs, it is better to have the crosssectional area of the horizontal segment of the U-tube to be much larger than the vertical segments. Hence, more energy is dissipated by the movement of

343

344

CHAPTER 5 Design procedures for tall buildings

the liquid in the horizontal segment and the vertical segments perform as a reservoir for the moving liquid (Sadek et al., 1998). Tube width ðbTLCD Þ. Given the tube width to liquid length ratio, rTLCD , and the liquid length, LTLCD , the tube width can be determined as: bTLCD 5 LTLCD rTLCD

•

(5.155)

MTLCDs   Liquid column length of each group LTLCD;j . Once the tuning frequency, fd;j , is determined from Eq. (5.151) for each group of MTLCDs, the liquid length, LTLDC;j , for ith group can be computed by: 2g 2 ω2 fd;j

(5.156)

μm PNd ρF j51 LTLCD;j

(5.157)

LTLCD;j 5

  Cross-sectional area of the tube ATLCD;j . Considering the same tube cross-sectional area, ATLCD;j 5 ATLCD , for each group of TLCDs, given the total mass ratio, μ (Step 7.1, Section 5.2.2.8.1) and liquid length, LTLCD;i , the following relation gives: ATLCD 5

  Tube width bLTCD;j . Considering rTLCD as the practical tube width to liquid length ratio, identical for all the groups of TLCDs, the tube width, bLTCD;j , associated with jth group can be determined as: bLTCD;j 5 LTLCD;j rTLCD

(5.158)

TLCD equivalent mechanical properties. Given the dimension of the TLCD tube, it is possible to determine the relative equivalent TMD mechanical properties (mass (md;eq ), stiffness (kd;eq ), and damping coefficient (cd;eq )). Simplified expressions are proposed by Chang and Qu (1998) and Chang et al. (1998) as follows: md;eq 5 ρF ATLCD LTLCD

(5.159)

kd;eq 5 2ρF ATLCD g

(5.160)

cd;eq 5

1 ρ ATLCD LTLCD ηd ju_ F j 2 F

(5.161)

where ηd is the head loss coefficient and u_ F is the liquid level velocity, which depends on the loading intensity. For a given loading intensity, iteration is necessary to define the velocity of the building
TLCD system (Chang and Qu, 1998).

5.2.2.9 Step 8: Update building model and perform analyses Given all the structural and damping properties of the building structure and mass damper, the designer is able to, first, update the mathematical model of the structure built in Step 3 (Section 5.2.2.4) and, then, explicitly model the damping

5.2 Passive Damping Systems

system with the use of commercial software packages. In this step, the hysteretic properties of structural elements may be accounted in the model. Depending on the type of the damping system chosen (TMD, TLD, TLCD), the corresponding mechanical parameters determined in Step 7 (Section 5.2.2.8) can be adopted. Eventually, the designer should conduct a time-history dynamic analysis, under wind or earthquake excitations, of the combined system of building structure and mass damper to evaluate the effectiveness of the damping system in mitigating vibrations and reducing desirable responses of the building (Tuan and Shang, 2014). The responses of interest, for example, lateral displacement, acceleration, and base shear, may be extracted from the history analysis. Note: it is recommended to use wind tunnel test results and/or computational fluid dynamics simulation to determine the wind loads acting on the building structure (Lee and Ng, 2010; Tuan and Shang, 2014). In the following, some recommendations for the modeling of the mass damping devices are given. •

•

•

TMD modeling The model can be built as a 3D structure, while the TMDs, depending on the installation position (e.g., top), can be modeled using link elements for springs and dampers (Farghaly and Ahmed, 2012). The mass, stiffness, and damping properties obtained in Step 7.2 (Section 5.2.2.8.2) can all be assigned to the link elements. According to CSI Knowledge Base (CSI, 2017), a TMD may be modeled using a spring
mass system with damping in SAP2000 (CSI, 2016b) or ETABS (CSI, 2016b). In this case, some useful guidelines are provided as follows:
Spring: Modeling a linear two-joint link with an assigned stiffness, kd , (spring) property in which one joint is attached to the structure, and the other one is free. The kd is obtained in Step 7.2 (Section 5.2.2.8.2)
Mass: Assigning TMD mass (md ) from in Step 7.2 (Section 5.2.2.8.2) to the free joint.
Damping: If SAP2000 (CSI, 2016b) is used, linear damping is included directly in the linear link property, while nonlinear damping is modeled using a viscous damping link object in parallel with the linear link. Using ETABS (CSI, 2016a), whether the system is linear or not, the damping objects are modeled in parallel. For the damping of TMD the value of cd , calculated in Step 7.2 (Section 5.2.2.8.2), should be used. TLD modeling In order to model a liquid damper because of the nonlinear behavior of the liquid, the TMD analogy according to the instructions given in the previous section can be utilized. The equivalent mass, stiffness, and damping coefficient of TLD systems are previously determined in Step 7.3 (Section 5.2.2.8.3). TLCD modeling Similar to the TLD modeling, the same approach can be used to model TLCD systems using the equivalent mass, stiffness, and damping obtained in

345

346

CHAPTER 5 Design procedures for tall buildings

Step 7.4 (Section 5.2.2.8.4). Sadek et al. (1998) state that the analysis results for SDOF systems equipped with TMDs and TLCDs give similar reductions in their responses for identical mass ratios. Therefore, TLCDs may be modeled as equivalent TMDs using commercial software. Lee et al. (2012) used the TMD analogy by involving equivalent supplemental viscous damping of the inherent nonlinear supplemental damping term of a TLCD.

5.2.2.10 Step 9: Check response acceptability The results obtained from the time-history analysis under design wind/earthquakes for the explicit model considered should be used to check all elements of the structure. For structures equipped with mass damping systems, strength requirements should be satisfied using the demands obtained from the nonlinear RHA under the DE (ASCE, 2017a; NEHRP, 2015). In addition, the following criteria should be satisfied: • •

•

Drift criteria. The same consideration for distributed dampers applied for this device category. Refer to Section 5.2.1.9.2 for further details. Acceleration criteria. Acceleration is considered the governing factor for human discomfort (see Chapter 3 for more details). For TMDs such as those in the John Hancock Tower in Boston or the Citicorp Center in New York, whenever the horizontal acceleration exceeds 0.003 g for two consecutive cycles, the system is automatically activated (Moon, 2010). Refer to Section 5.2.1.9.3 for further details. Damper criteria. The mass damping system response shall satisfy all design requirements defined during the procedure. Moreover, the demand in the system shall be sent to the mass damper consultants for the mechanical verification of the system. Regarding code requirements, the general consideration provided by NEHRP (2015) (for distributed dampers, Section 5.2.1.9.4) can be applied for this device category as well.

Note that in case the response results obtained in Step 8 (Section 5.2.2.9) may not satisfy the design criteria, the present design procedure is potentially iterative. In this case, the designer must refer back to Step 5 (Section 5.2.2.6) in order to reselect a new effective total damping and repeat subsequent steps anew. This may lead to an update of the damper parameters such that the considered criteria are satisfied.

5.2.2.11 Step 10: Quality assurance and experimental evaluation After the design of the building and relative damping components a quality control plan shall be written by the design professional to verify the following items (NEHRP, 2015): • •

Manufacturing process Preinstallation tests (e.g., prototype and production tests) to validate the design mass damping properties and force
velocity
displacement relationships

5.3 Isolation Systems

• •

Commissioning and system tuning Inspection/maintenance procedures A detailed discussion on all these aspects is given in Chapter 7.

5.3 ISOLATION SYSTEMS 5.3.1 STEP-BY-STEP PROCEDURE FOR BASE ISOLATION The basic idea of seismic isolation is that the superstructure (that is above the isolation system) is decoupled from the horizontal earthquake ground movement (Section 4.2 (Chapter 4)). This is achieved by having a stiff superstructure and a flexible isolation interface. In this way, most of the deformation demand tends to concentrate on the isolation system and therefore the superstructure tends to behave as an SDOF system (mass of the superstructure and stiffness of the isolation system). This behavior produces less overturning moments, less accelerations demand (with less content damages and less earthquake perception), less interstory drifts (with less structural and nonstructural damages), but greater deformation demand, which is mostly concentrated at the isolation interface. This greater deformation level produces two undesired effects: first, it requires a gap at the perimeter in order to avoid collision with the adjacent structures (as well as for incoming building utilities) and second, it lowers the compression capacity of the isolators due to P-delta effects. Therefore, the increase in displacement needs to be controlled and for this scope, adding supplemental damping to the isolation system could be a solution. The addition of damping will lower the isolation system’ displacement demand and the amount of energy that excites the superstructure by further improving the efficiency of the isolation system (see Figs. 3.26 and 3.27 (Chapter 3)).

5.3.1.1 Design procedures literature review The procedures for design of base-isolated structures under seismic excitations are mainly developed based on the provisions and methods addressed in Section 5.1.1. ASCE 7-16 (ASCE, 2017a) and FEMA P-1050-1 (NEHRP, 2015) provide a design approach for new seismically isolated structures in which structural and isolation systems are designed according to the MCER demand. They allow the adoption of the equivalent linear analysis (not permitted for tall buildings), RSA or NLTHAs. Similarly, for the retrofit of existing structures, ASCE 41-13 (ASCE, 2013) provides recommendations as shown in Section 5.5. Similar to US design codes for the design of base-isolated structures, Eurocode 8 (CEN, 2004) allows the use of equivalent linear analysis, RSA, or NLTHAs. The Italian code (NTC, 2008), based on the Eurocode 8 (Cen, 2004), introduces a collapse limit (MCER-type earthquake) for designing isolators in parallel to a DE level for designing the superstructure. The Chinese code GB 50011 (GB50011, 2010) also

347

348

CHAPTER 5 Design procedures for tall buildings

addresses the design of multistory (RC) structures with isolation stories, in which the structure is modeled as shear-type system to be analyzed using time-history analysis. The Japanese code (JSSI Design Guideline and Manual (JSSI, 2010)) gives the basic principles of designing seismically isolated buildings by using time-history analysis to examine their safety under earthquake excitation. The design of both new and retrofitted buildings is in the scope of Japanese code. Among all the available procedures in most of the cases, equivalent linear analysis techniques have been used for the simplified design of base-isolated buildings. This has been proposed by both codes (e.g., FEMA (1997); ICBO (1997); NEHRP (1997); AASHTO (2002); ICC (2003); FEMA (2003); CEN (2004); ASCE (2005); NTC (2008); AASHTO (2010); ASCE (2010); ASCE (2013); ASCE (2017a)) and researchers (e.g., Hwang and Chiou (1996); Kwan and Billington (2003); Blandon and Priestley (2005); Jara and Casas (2006); Dicleli and Buddaram (2007); Jara et al. (2012); Liu et al. (2014)). Several research works discuss procedures based on standard design codes, which are briefly revised as follows: •

•

•

•

•

•

Naeim and Kelly (1999) presented a step-by-step procedure for base isolation design, which is compliant with requirements of UBC (1997). This procedure consisted of a preliminary and a final design phase that include the steps for analyses and verifications of the base isolation system. A detailed example using the proposed procedure was provided. Higashino and Okamoto (2006) in their textbook presented a design flowchart based on the Japanese JSSI Design Guideline and Manual (JSSI, 2000). The design example of a 10-story RC structure with the use of this step-by-step procedure was presented. Christopoulos and Filiatrault (2006) proposed two types of design methods for base isolation systems of buildings: static and dynamic analyses. The methods were mostly based on those provided in ICC (2003) and FEMA (2003). Wen and Baifeng (2008) proposed a two-step design method based on Chinese code (GB50011, 2010). The first step was devoted to the estimation of isolation layer from few basic structural data. The second step was adopted to address a detailed design of superstructure, foundation, and base isolation devices. Kani et al. (2010) represented a design flowchart by following the approach presented in MRIT (2000), with equivalent linear and time-history analysis methods. They outlined that isolated buildings higher than 60 m should be designed using time-history analysis method. A 7-story base-isolated RC building was designed using the proposed design procedure. Feng et al. (2012) proposed a preliminary two-stage design procedure, called CW2012, for seismically isolated buildings, based on international seismic isolation codes (MRIT (2000); ASCE (2005); NTC (2008); ICC (2009); GB50011 (2010)). This procedure was mainly based on Japanese 2000 code (MRIT (2000)) and recommends a time-history analysis method, proposed by

5.3 Isolation Systems

•

JSSI (2010). It also addresses the use of the equivalent linear analysis method with highlighted limitations. Becker et al. (2015) reported that Japanese design code (e.g., SSI, 2000) has clearly outlined the procedures for designing isolated high-rise buildings. Using isolation systems in high-rise buildings in Japan could demonstrate that is practical also under US code (ASCE, 2010). However, US requirements are considerably more stringent than the Japanese ones (JSSI, 2000).

In addition to code prescriptive design procedures, several researches have proposed alternative ones. A brief review of the most important is the following: •

•

•

•

Shinozaki et al. (2004) addressed useful design issues of tall building equipped with base isolation systems. They reported the high seismic performance in the design of two built base-isolated tall buildings (over 60 m high). Pan et al. (2005) presented basic design procedures based on common Japanese design practice. In addition, some design issues of base-isolated buildings are reviewed: variation in base isolation material properties, how to apply for high-rise buildings, influences of vertical ground motions, and response under near-fault ground motions. Lee and Liang (2012) proposed a comprehensive review of seismic isolation design principles in their textbook chapter based on SDOF and MDOF models. Islam et al. (2011, 2012) developed a simplified step-by-step design procedure. This method was separately proposed for two isolator materials: lead
rubber bearings (LRBs) and high-damping rubber bearings (HDRBs). The proposed design procedures were proposed to be included in Bangladesh National Building Code (BNBC, 1993).

5.3.1.2 Step 1: Building and site categorization In the first step, tall building categorization (e.g., the building risk category and relative occupancy importance factor) and relative site categorization (e.g., spectral response acceleration and response spectrum) should be specified. For more details about above-mentioned information, the reader is encouraged to refer to Step 1 of Section 5.2.1.1. However, for seismically isolated structures, the seismic importance factor shall be taken as 1.0 regardless of risk category assignment.

5.3.1.3 Step 2: Select the structural system(s) A proper selection of a seismic load-resisting structural system(s) is desired. To this end, Section 5.2.1.2 provides adequate details for designers based on ASCE 7-16 (ASCE, 2017a). Structural system should minimize the axial loads induced by the overturning moment to avoid tension/uplift or excessive compression forces on the isolators, especially on taller buildings. This is more easily achieved

349

CHAPTER 5 Design procedures for tall buildings

with structural systems based on moment frames rather than core wall because they typically have a greater level arm. Avoiding tension forces (or uplifts) on the isolators’ units usually determines the limit of whether or not seismically isolating a tall building is feasible. In addition, the aspect ratio H/L (i.e., the ratio between height H and width L of the building) has several limitations as it can be found in the literature (Li and Wu, 2006): •

Site soil condition. Hard sites permit higher aspect ratio limits than softer sites under every seismic intensity (Fig. 5.23). Hence, base-isolated tall buildings (as well as not isolated ones) are preferred to be built on hard sites, and it is dangerous to build them on soft sites with potential strong ground motions.

(A) 18 16 Hard site

Limit ratio of H/B

14 12 10 8 6

Medium site

4

Soft site

2 0 0.2

0.3

0.7 0.6 0.4 0.5 Peak acceleration of ground (g)

0.8

(B) 18 16

Hard site

14 Limit ratio of H/B

350

12 10 8 6 Medium site

4

Soft site 2 0

1

1.2

1.4 1.6 1.8 2 Period of isolated structure (s)

2.2

2.4

FIGURE 5.23 Aspect ratio estimated limits in rubber-type base-isolated buildings against (A) peak ground acceleration and (B) period of isolated structure (Li and Wu, 2006).

5.3 Isolation Systems

• •

Ground motion intensity. Higher ground acceleration generally leads to a decrease in the aspect ratio of isolated buildings (Fig. 5.23A). Period of isolated system. Isolated buildings with longer (fundamental) periods may exhibit higher aspect ratio limits (see Fig. 5.23B); this relationship is more dominant when the site is harder.

5.3.1.4 Step 3: Target vibration period and total damping ratio selection The appropriate selection of the target vibration period TM for MCER is based on the achievement of a desired level of acceleration and displacements based on spectrum demand. Based on the target vibration period and the mass m of the superstructure, the target global stiffness of the isolation system can be preliminary estimated based on the equations for an SDOF system: KM 5 4π2

m 2 TM

(5.162)

Note that this estimation will be more accurate if the superstructure is very stiff, which cannot be the case in taller buildings. The target total damping ratio (ζ TM ) of the isolated building should be selected by the designer, so that a desired reduction in a target response (e.g., base shear, superstructure accelerations or isolation system displacement demand) can be achieved. Usually based on the selected base isolation system (Step 4, Section 5.3.1.5) the relative added damping can be estimated according to experimental tests or known information. For example, for isolators with high-damping rubber material Naeim and Kelly (1999) suggested to using a 10
12% damping ratio. According to ASCE 41-13 (ASCE, 2013), the effective damping created by elastomeric isolators is typically less than 7% for shear strains in the range of 0
2, while with LRBs and frictional devices damping ratios larger than 15% can be usually be obtained.

5.3.1.5 Step 4: Isolation type and distribution In this stage, it is useful to decide which kind of isolation (bearing) systems is suitable for the building under study to achieve the desired target vibration period and damping ratio. According to NEHRP (2009) an acceptable isolation system should have the following characteristics: • • • • •

Stable for maximum earthquake displacements Increasing resistance with increasing displacement Limited degradation under repeated cycles of earthquake load Minimum restoring force requirement Well-established and repeatable engineering properties (effective stiffness and damping).

5.3.1.5.1 Determine type of isolation system The most common type of seismic isolation systems used in US and Japanese buildings are (Higashino and Okamoto, 2006): HDRB, LRB, sliding bearing

351

352

CHAPTER 5 Design procedures for tall buildings

systems, such as the friction pendulum system, or some combination of elastomeric and sliding isolators. If several types of isolators are combined, one type should be dominant (Pan et al., 2005). NRB and LRB were often used because of the lower cost compared to HDRB types and the difficulty of achieving large isolated periods due to the increased lateral stiffness of HDRBs. The sliding bearings are frequently used in combination with NRB, LRB, or HDRB to decrease the stiffness of the isolation layer for the large displacement domain (Pan et al., 2005) and be able to achieve large target vibration periods. See Chapter 4 (Section 4.2) for a detailed discussion on base isolation devices. Seismic base isolation solutions become usually less effective for high-rise buildings due to long periods of the fixed-base structure configuration and also because wind forces may govern the design of these types of structures and isolation systems are generally not activated under wind loads. Therefore, it may be recommended to incorporate energy dissipation devices (e.g., viscous dampers) at the isolation layer in such buildings (Pan et al., 2005; Becker et al., 2015). This is so called hybrid isolation system. The advantages of using dampers in conjunction with isolators could be: • • •

Reduction in relative displacement demand in isolation layer (Hall and Ryan, 2000; Pan et al., 2005; Wolff et al., 2014) Reduction in story drifts (or even floor acceleration) when viscous dampers are used (Hall and Ryan, 2000) Reducing impact of near-fault ground motion (Ariga et al., 2006; Providakis, 2008)

However, using such dampers, with too high damping, may be detrimental when the isolated building is excited by far-field earthquakes (Hall and Ryan, 2000; Providakis, 2008; Wolff et al., 2014). In Japan, the most common hybrid isolation solutions utilized for high-rise building are (Pan et al., 2005): • • • •

NRBs combined with lead dampers NRBs combined with steel coiled dampers LRBs combined with steel, lead, or oil dampers HDRBs combined with steel, lead, or oil dampers

Applications of hysteretic dampers (at the isolation level) were carried out by Okamoto et al. (2002) and Pan et al. (2005) that recommended to adopt a total yield force of hysteretic-type dampers (e.g., steel or lead dampers) between about 3% and 5% of the total weight of superstructure. An additional study carried out by Takewaki (2008) showed the influence of viscous (from viscous dampers) and friction (from friction-type bearing) damping in dissipating energy (to attain a given target drift in base isolation) for different building heights. It was demonstrated that for taller base-isolated buildings (40-story), the role of friction damping (i.e., friction-type isolator) is predominant than in lower rise isolated buildings; moreover, the overall damping ratio is less. Despite these studies an

5.3 Isolation Systems

alternative solution was proposed by Komuro et al. (2005) that introduce the hybrid TAISEI Shake Suppression system in which only a combination of sliding and rubber bearings are utilized without requiring additional dampers. Komuro et al. (2005) demonstrated that this system can lead to similar strong seismic performance for high-rise buildings of a base isolation that has additional dampers (such as steel or lead dampers) without a large increase in the construction costs as in the case of the utilization of additional dampers. Similarly in the United States, the following typical combinations are utilized (Wolff et al., 2014): • • • •

Elastomeric bearings with nonlinear viscous dampers HDRBs combined with nonlinear viscous dampers Single friction pendulum bearings with nonlinear viscous dampers Triple/quintuple friction pendulum bearings with linear viscous dampers

5.3.1.5.2 Determine location and distribution of isolation system Having defined the type of isolation devices, target period (and therefore isolation system stiffness) and target damping, the next step is to determine the distribution of the isolators in the building. Isolation systems do not necessarily need to be at the base of a structure. They can be at any level; however, only the elements above the isolation system are decoupled from the horizontal earthquake ground movement and are therefore seismically protected by the isolation system. For this reason, codes require that structures below the isolation level are designed for the elastic seismic demand, which in some cases produces substructure elements larger in size than usual. Another important aspect is that vertical circulation systems need to be detailed to accommodate movement at the isolation interface (see Chapter 6 for further details). In general, isolators are located in plan based on the vertical load paths. Usually, there is a need for at least one isolator under each column and at least two under each wall (depending on wall length). Transfer systems may be used with caution in order to decrease the number of isolator required and therefore achieve the target vibration period defined in Step 3 (Section 5.3.1.4) more easily. Moreover, friction bearing-type isolators should have larger tributary loads to activate the system; this is why they are preferably placed in interior locations. The isolator stiffness distribution is chosen in order to minimize eccentricity with superstructure mass. Moreover, isolators that provide higher dissipation are usually located on the perimeter of the plan to increase energy dissipation and to improve torsional stiffness (see Fig. 5.24 for an isolator plan distribution example for Nunoa Capital Building in Chile and relative case study in Section 8.3.1 (Chapter 8)). Note: it is worth mentioning that the isolated plane should have a robust and relatively stiff diaphragm to transmit the lateral loads in a relatively uniform fashion.

353

354

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.24 Plan distribution example for Nunoa Capital Building, Santiago de Chile. Courtesy of Rene Lagos Engineers.

5.3.1.6 Step 5: Preliminary structural analysis Preliminary structural analysis should be performed in order to obtain an estimation of the isolation system displacement and force demand for the MCER, so that a preliminary design of the isolators can be made. According to US design codes, ASCE 7-16 (ASCE, 2017a), ASCE 41-17 (ASCE, 2017b), and FEMA P-1050-1 (NEHRP, 2015), seismically isolated structures shall be designed using dynamic procedures (response spectrum and response history). ELF procedure is also allowed only when the following requirements are met: • • • • • •

For building in site classes A, B, C, and D. Period of isolated structure at maximum displacement less than or equal to 5.0 second and greater than 3 times the period of the fixed base structure Height of structural above isolation system less than 19.8 m. Effective damping of isolation system at maximum displacement is less than 30%. No structural irregularity of the building above the isolation level. The isolation system shall meet the following requirements: • The effective stiffness at maximum displacement is greater than one-third of the effective stiffness at 20% of the maximum displacement. • Have restoring force. • The isolation system does not limit the maximum earthquake displacement.

5.3 Isolation Systems

Despite the fact that it is not allowed for all types of buildings, the ELF method provides a good starting point for preliminary isolation system demand estimation (displacements and forces) since both dynamic procedures (response spectrum and nonlinear time history) require minimum displacement and shear demands based on the results of the ELF procedure. Furthermore, ELF allows a better understanding of the force flow as the CQC and SRSS methods use square root of squared results. For these reasons, in the following, the basic principles of ELF are briefly reviewed. The first step consists in calculating the maximum displacement demand and corresponding effective vibration period as follows: •

Maximum displacement: DM 5

•

gSM1 TM 4π2 Bζ;TM

(5.163)

Effective period at maximum displacement: sffiffiffiffiffiffiffiffi W TM 5 2π kM g

(5.164)

where Bζ;TM is the coefficient related to the total effective damping (isolation plus intrinsic) at maximum displacement (ζ TM ) for the isolated modes, per Table 5.18. According to ASCE 7-16 (ASCE, 2017a) the damping contribution from the isolator is computed based on the upper- and lower-bound force
deflection behavior of individual isolator units, as follows: P

ζ dM 5

EM 2πkM D2M

(5.165)

P where EM is the total energy dissipated in the isolation system during a full cycle response at maximum displacement. kM is the effective stiffness of the isolation system as selected in Step 4 (Section 5.3.1.5). According to ASCE 7-16 (ASCE, 2017a) the effective stiffness shall be based on the upperand lower-bound force
deflection behavior of individual isolator units, as follows: kM 5

P 

 P  2 1 F  FM 1 M 2DM

(5.166)

P  1  P  2  where FM and FM are the sum of the force absolute values for all isolators at positive and negative maximum displacements, respectively. The total displacement of the elements in the isolation system needs to take into consideration the actual and accidental torsion, based on the spatial stiffness distribution and the critical location of the eccentric mass. The maximum total

355

356

CHAPTER 5 Design procedures for tall buildings

displacements for uniform spatial distribution of lateral stiffness should be calculated using the following equation:   y 12e $ 1:15DM DTM 5 DM 1 1 2 2 PT L1 1 L22

(5.167)

where y is the distance between the center of rigidity of the isolation system and of the element under consideration, measured perpendicular to the loading direction; e is the eccentricity, measured as the distance between the center of mass of the structure above the isolation interface and the isolation system itself. Additionally, an accidental eccentricity of 5% of the longest plan dimension, measured perpendicular to the loading direction, shall be included; L1 is the structure’s longest plan dimension; L2 is the structure’s shortest plan dimension, measured perpendicular to L1 ; and PT is the ratio of translation to torsional period of the isolation system as determined from dynamic analysis or from the following equation (ASCE, 2017a): 1 PT 5 rl

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi PNd  2 2 i51 xi 1 yi $1 Nd

(5.168)

where xi and yi are the two horizontal axes distances of each isolator unit i from the center of mass; Nd is the number of isolator units; rl is the radius of gyration qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  ffi of the isolation system as L21 1 L22 =12. The second step consists in calculating the minimum design lateral forces for the elements above and below the isolation system. For all the structural elements below the isolation system, the maximum lateral force (VMb ) can be determined as follows: VMb 5 kM DM

(5.169)

For the structural elements above the isolation system, the design shear force, VMa , shall be calculated as follows: VMa 5

Vst RI

(5.170)

where RI is a factor equal to three-eighths the response modification factor of the structural system above isolation (as per Table 5.16) with a value between 1 and 2 and Vst is the total unreduced seismic design force for the elements above the isolation level, determined as follows:  ð122:5ζ TM Þ Ws Vst 5 VMb W

(5.171)

5.3 Isolation Systems

where W and Ws are the effective seismic weights of the structure above the isolation level with and without the effective weight of the base level. In case the distance between the top of the isolator and the underside of the base level floor 0.9 m the two values are equal. Note the exponential term is greater than  1 2 2:5ζ TM shall be replaced by 1 2 3:5ζ TM in case the isolation system hysteretic behavior has an abrupt transition from the preyield and postyield (or preslip to postslip) behavior. ASCE (2017a) provides three limits on the value of the shear force above the isolation level as follows: •

• •

Lateral shear force required by code for a fixed-based structure (Eq. 5.44) of the same seismic weight Ws and isolated period TM using upper-bound properties of the isolation system Factored design wind load base shear (Section 5.2.1.1.4) Vst calculated as per Eq. (5.171) with VMb set as the force to fully activate the isolation system using the greatest between the upper bound properties, or: • 1.5 times the nominal properties for the yield level of a softening system • The ultimate capacity of a sacrificial wind-restraint system • The breakaway frictional force of a sliding system • The force at zero displacement for the sliding system following a complete cycle of motion at DM .

Subsequently, the shear force above the isolation level should be distributed vertically as follows (ASCE, 2017a): VMa wx hkx fx 5 Pn k i52 wi hi

(5.172)

where fx portion of VMa that is assigned to level x; wx portion of W that is located at or assigned to level x; hx height above the base of level x; and k 5 14ζ MT TMb , where TMa is the period of the structure above the isolation level computed assumed a fixed-based system. Having determined the distributed forces in the structure above the isolation level, the axial forces on each isolator due to both static loads and seismic overturning moments should be estimated.

5.3.1.7 Step 6: Isolation system preliminary design The design of the isolation system is an iterative process. For example, leadplug and friction pendulum isolation bearings can be generally modeled using a bilinear hysteretic model (see Section 4.2.1 (Chapter 4) and Naeim and Kelly (1999)). Three main parameters are utilized to describe this behavior: elastic stiffness, k0 ; postyield stiffness, k0p ; and characteristic strength f0 . Herein, k0 is estimated from test data for elastomeric bearing or as a multiple of k0p for

357

358

CHAPTER 5 Design procedures for tall buildings

lead-plug and friction pendulum bearings. The characteristic strength for a lead-plug bearing is a function of the yield stress, whereas for a friction pendulum bearing, it is a function of the friction coefficient and the vertical load. These values are calculated for the (effective) target vibration periods ((for each isolator unit) TM ) of Step 3 (Section 5.3.1.4), displacement demands (DM ) of Step 5 (Section 5.3.1.6), (effective) target stiffnesses (kM ) of Step 3 (Section 5.3.1.4), and (effective) target total damping (ζ TM ) of Step 3 (Section 5.3.1.4). Based on these parameters, the characteristic strength can be initially estimated by neglecting the yield displacement (Dy 5 0) (Naeim and Kelly, 1999): f0M 5

EDM π 5 kM DM ζ TM 2 4DM

(5.173)

where the index M stands for MCER demand and EDM is the area under the hysteresis loop, computed as: WDM 5 2πkM D2M ζ TM

(5.174)

The postyield stiffness can then be derived from the following expression: k0pM 5 k0M 2 f0M =DM

(5.175)

The elastic stiffness can be assumed as a function of the postyield stiffness and the bearing type. For example, Naeim and Kelly (1999) proposed the following relationships: k0M 5 10k0pM for lead-plug bearing K1M 5 100k0pM for friction pendulum bearing

(5.176)

Having determined the elastic and the postyield stiffnesses the first estimation of the yield displacement can be obtained from the following expression (Naeim and Kelly, 1999): DyM 5

f0M k0M 2 k0pM

(5.177)

Afterward the characteristics strength can be updated based on the yield displacement as follows: 2 2π Wg T2πM D2M ζ TM   f0M 5 4 DM 2 DyM

(5.178)

Finally, the postyield stiffness can be updated as well using Eq. (5.175). It is worth mentioning that the value of k0M may vary over a wide range and while it has no influence on the effective stiffness, it significantly influences the supplemental damping (Naeim and Kelly, 1999).

5.3 Isolation Systems

For the specific case of friction pendulum bearings, simple expressions to predetermine design parameters (radius of concave plate r, friction coefficient rμ , and the initial stiffness k0 ) are listed, respectively, as (Naeim and Kelly, 1999): rM 5

2 gTM ð2πÞ2

 πζ TM DM  rM 2 2 πζ TM  rμM 1 k0M 5 W 1 DM rM

rμM 5 

(5.179) (5.180)

(5.181)

Generally speaking, the size of a bearing depends upon the amount of vertical load that should be sustained by it, the target stiffness/damping, and the maximum amount of lateral earthquake displacement that must be accommodated (FEMA 751 (NEHRP, 2012)).

5.3.1.8 Step 7: Update building model and final analyses After the preliminary design of the isolators is completed, the model must be updated and dynamic analysis should be performed. ASCE (2017a) requires to utilize a mathematical model that includes the isolation system, the seismic forceresisting system, and all the other structural elements that influence the dynamic behavior of the structure. The design of both isolation system and structure above shall be performed separately for upper- and lower-bound properties of the isolation system. The structure above the isolation system is permitted to be modeled as a linearly elastic if the seismic force-resisting system is essentially elastic under a lateral force not less than 100% VMa (from Eq. 5.170). Two alternatives dynamic analysis methods are allowed by ASCE (2017a): RSA and NLTHA. Each of them has its own advantages and disadvantages. RSA could be more suitable to design the structure, but it considers the response reduction due to additional damping indirectly by dividing the spectrum demand by the Bζ factor (Section 5.2.1.4.1) and may not be suitable for friction-based devices. Instead, in NLTHA the energy dissipation is directly modeled through hysteretic behavior. Regardless of the dynamic analysis procedure chosen, according to ASCE 716 (ASCE, 2017a), the elements above the isolation level shall be designed according to the requirements for no-isolated structure with a seismic force computed as per Eq. (5.170). Additional requirements depending on the analysis carried out are the following: • •

For RSA the lateral force for the structural elements above the isolation system shall not be less than VMa (from Eq. 5.170) For RHA of regular structures, the design lateral force, for the structural elements above the isolation system, VMa (from Eq. 5.170) shall not be taken as less than 80% of the VMb (from Eq. 5.169) computed from an ELF analysis.

359

360

CHAPTER 5 Design procedures for tall buildings

•

Moreover, the limits on the base shear, VMa , as specified in Step 5 (Section 5.3.1.6) for ELF analysis shall be verified. For RHA of irregular structures, the design lateral force, for the structural elements above the isolation system, VMa (from Eq. 5.170) shall not be taken as less than 100% of the VMb (from Eq. 5.169) computed from an ELF analysis. Moreover, the limits on the base shear, VMa as specified in Step 5 (Section 5.3.1.6) for ELF analysis shall be verified.

5.3.1.8.1 Response-spectrum analysis For RSA the modal damping, for the fundamental mode in the main direction, shall be not greater than the isolation system effective damping 30% of critical. Instead, for higher modes, appropriate damping values shall be selected for the structure assuming a fixed base. To determine the total maximum displacement (under MCER), simultaneous excitation forces should be considered with 100% of the ground motion in the desired direction and 30% of the ground motion in the perpendicular direction (ASCE, 2017a). The total displacement is then calculated as the vector sum of the two orthogonal displacements.

5.3.1.8.2 Nonlinear time-history analysis For nonlinear time history a set of not less than seven pairs of horizontal acceleration components shall be considered with characteristics similar to those of the control MCER (ASCE, 2017a; PEER, 2017). Amplitude and spectral matching is allowed to scale the ground motion (see Section 5.2.1.4.2 for more details). Each pair of ground motion shall be applied simultaneously to the model considering the most disadvantageous condition of eccentric mass (in addition to an accidental mass eccentricity equal to 5% of the diaphragm dimension). The maximum displacement of the isolation system shall be calculated as the vector sum of the two orthogonal directions per each time step and the average of each response parameter shall be used for design.

5.3.1.8.3 Isolation system modeling The isolation system shall be modeled in order to accurately represent its behavior. The isolators are modeled with their nonlinear force
deflection characteristics obtained from Step 6 (Section 5.3.1.7). The characteristics of the isolation system must be justified through testing (Section 7.2.3 (Chapter 7)) and must consider the nonlinear properties of the system itself. In literature, the following recommendations for isolator modeling can be found: •

Single concave friction pendulum. FEMA 751 (NEHRP, 2012) recommends modeling this category of devices with a nonlinear model. This model usually requires the user to input the initial stiffness (effective stiffness before the slider displaces on the concave plate), the friction coefficient, and the radius

5.3 Isolation Systems

•

•

•

•

of the concave plate (Eqs. 5.179
5.181). Furthermore, it also requires the input of three additional sliding friction parameters (NEHRP, 2012): 1. “Fast speed” sliding coefficient of friction. This value represents velocities more than approximately 2.5
5.1 cm/s. Typically, during strong ground motions, sliding velocities are well in excess of 5.1 cm/s and isolation system response is dominated by “fast speed” sliding friction properties. 2. “Slow speed” sliding coefficient of friction. 3. “Rate” parameter that essentially governs the transition between slow and fast speed properties (e.g., when the bearing reaches peak displacement and reverses direction of travel). Double/triple/quintuple friction pendulum. It requires more sophisticated software (e.g., 3D-BASIS (Nagarajaiah et al., 1991; Reinhorn et al., 1994; Tsopelas et al., 1994a,b; SAP2000 (CSI, 2016b)). Indeed, different friction coefficients for all the surfaces of top and bottom concave plates need to be defined. The different modeling possibilities for multiple friction pendulum are discussed in detail by Lee and Constantinou (2015). The isolation system should be modeled to account for the spatial distribution of isolator units which is determined in Step 4 (Section 5.3.1.5) (ASCE, 2017a). The effects of vertical load, bilateral load, and/or the variability of isolation system properties (e.g., due to rate of loading) on the force
deflection characteristics of the isolation system must be accounted for in the modeling (ASCE, 2017a). Avoid damping leakage (Sarlis and Constantinou, 2010) that is described as “the undesirable scenario where the user-assigned modal damping in an isolated model is effectual for the modes corresponding to the isolation system, thus ‘leaking’ into the isolation system and ultimately resulting in underestimation of the actual response. This phenomenon can be easily controlled in the FNA because the method gives the capability to the user to assign global values of damping for all the modes (e.g., by Rayleigh, constant, and linear damping) while at the same time damping values can be manually overridden for a number of modes, a feature that is not available in the nonlinear DI time-history analysis method” (Oikonomou et al., 2016).

It is important to understand that variabilities in isolation properties may occur due to several factors, such as manufacturing tolerances, effects of heating during cyclic motion, effects of aging, contamination, ambient temperature, and history of loading. To take into consideration these variabilities, modification factors are utilized (Table 5.27 (NEHRP, 2015)). These factors multiplied by the nominal properties (as shown in Section 5.1.3.3.3) produce upper and lower bound force
deflection loops of the isolation system, as per Eqs. (5.4) and (5.5). Then, dynamic analyses are performed and the resulting demands are enveloped for design purposes. It is worth mentioning that the most preferred method to establish property modification factors is to utilize rigorous qualification testing of

361

Table 5.27 Default Upper and Lower Bound Multipliers for Isolation Bearings (NEHRP, 2015) Variable Aging and Environmental Factors and Testing Factors

Sliding Bearing Unlubricated Interfaces μa or Qd b

Sliding Bearing Lubricated (liquid) Interfaces μa or Qd b

Plain Low Damping Elastomeric k

LRBc kd

LRB Qd

HDRBd kd

HDRB Qd

Aging λa Contamination λc Example upper bound λae;max Example lower bound λae;min All cyclic effects, upper All cyclic effects, lower Example upper bound λtest;max Example lower bound λtest;min λPM;max (1 1 (0.75 ( λae;max 21))  λtest;max λPM;min (1- (0.75 (1
λae;min ))  λtest;min Lambda factor for spec. tolerance λspec;max Lambda factor for spec. tolerance λspec;min Upper bound design property multiplier Lower bound design property multiplier Default upper bound design property multiplier Default lower bound design property multiplier

1.3 1.2 1.56 1 1.3 0.7 1.3 0.7 1.85

1.8 1.4 2.52 1 1.3 0.7 1.3 0.7 2.78

1.3 1 1.3 1 1.3 0.9 1.3 0.9 1.59

1.3 1 1.3 1 1.3 0.9 1.3 0.9 1.59

1 1 1 1 1.6 0.9 1.6 0.9 1.6

1.4 1 1.4 1 1.5 0.9 1.5 0.9 1.95

1.3 1 1.3 1 1.3 0.9 1.3 0.9 1.59

0.7

0.7

0.9

0.9

0.9

0.9

0.9

1.15

1.15

1.15

1.15

1.15

1.15

1.15

0.85

0.85

0.85

0.85

0.85

0.85

0.85

2.12

3.2

1.83

1.83

1.84

2.24

1.83

0.6

0.6

0.77

0.77

0.77

0.77

0.77

2.1

3.2

1.8

1.8

1.8

2.2

1.8

0.6

0.6

0.8

0.8

0.8

0.8

0.8

a

Friction coefficient in sliding bearings. Characteristic strength in sliding bearings. c Lead
rubber bearing. d High-damping rubber. b

5.3 Isolation Systems

materials and manufacturing methods by a qualified manufacturer (see Chapter 7 for more details about testing). In general, for most isolation systems, no modification factors for ambient temperature effects are considered, if they are located in a conditioned space, where the anticipated temperature varies between 21 C and 38 C (NEHRP, 2015). Moreover, thermal effects can be important for sliding bearings and lead
rubber bearings, while no consideration is needed for elastomeric bearings of either low or high-damping rubber (Constantinou et al., 2007). In addition to the specific recommendation for the modeling of isolation devices, the following considerations, for the analysis of structures equipped with isolation devices, are important: •

•

•

•

•

The P-delta effects which can be quite significant on the isolation system and adjacent elements of the structure shall be considered in modeling (NEHRP, 2012). The maximum displacement of the isolation system shall be calculated from the vectorial sum of the two orthogonal displacements computed at each time step. Appropriate selection and scaling of ground motions should be done by a ground motion expert having experienced with earthquake hazard conditions of the region, considering site conditions, earthquake magnitudes, fault distances, and source mechanisms that influence ground motion hazard at the building site (NEHRP, 2012). To account for the seismic load effects, the load combinations recommended in Section 5.2.1.1.4 can be considered for isolated structures (ASCE, 2017a; NEHRP, 2015). A nonlinear static analysis for an MCER displacement of the entire structural system, including the isolation system, is also allowed to demonstrate the isolation system stability and showing that the lateral and vertical stabilities are maintained (ASCE, 2013; NEHRP, 2015).

Lateral displacement due to wind loads over the depth of the isolation system should be limited to a value similar to that required for other stories of the superstructure.

5.3.1.8.4 Usual practice The usual practice is to perform only nonlinear time history, or nonlinear time history for isolation and global performance followed by ELF distribution of resulting story shear distribution. From the NLTHA, the following values are obtained: •

Maximum isolation system demands • Maximum displacements (DM , DTM ) • Maximum axial loads on isolators

363

364

CHAPTER 5 Design procedures for tall buildings

•

•

Effective properties • Effective stiffness of isolation system (kM ) • Effective vibration periods of the structure (TM ) POs • Maximum floor accelerations • Maximum interstory drifts

These values can be used to validate the preliminary base isolation design carried out with ELF. Uplifts on isolators should be avoided and maximum compression should be feasible to be applied in a laboratory test. Also, the gap between isolated and nonisolated structure must be greater than DTM (Fig. 5.25). Moreover, the response acceptance criteria in terms of admissible accelerations and interstory drifts should be met.

5.3.1.9 Step 8: Quality assurance An independent engineering team, with one or more individuals with a minimum of one registered design professional, should conduct a design review of the isolation system and related test programs. The review shall include at least the following criteria (ASCE, 2017a): • •

Site-specific spectra and ground motion histories Isolation preliminary design: device selection and determination of force and displacement levels

FIGURE 5.25 Gap (red perimeter) between isolated (light gray) and nonisolated (dark gray) structures. Courtesy of Rene Lagos Engineers.

5.4 Active, Semiactive, and Hybrid Systems

• • • •

Appropriate selection of property modification factors Prototype testing program Design of isolation units and structure Production testing program

Quality assurance programs are defined by the structural engineer, while prototype tests (Chapter 7) for each type of isolator must be performed by the isolator suppliers in order to confirm the effective stiffness and damping ratios assumed in the design as assumed in the previous steps. For further details, refer to Chapter 7 (Section 7.2.3).

5.4 ACTIVE, SEMIACTIVE, AND HYBRID SYSTEMS 5.4.1 LITERATURE REVIEW In Chapter 4 (Section 4.3), the main principles of active, semiactive, and hybrid damping systems were introduced. In the present section, the design procedures for buildings with the most common types for this category of dynamic modification device are reviewed. It is worth mentioning that no standard code is available for the design of active, semiactive, and hybrid control systems. Thus, the proposed design procedure is essentially based on the existing research works available in the literature. These methods (that are mainly for active-tuned mass damper (ATMD) and magnetorheological (MR) dampers) are briefly reviewed in the following: •

•

•

•

Chang et al. (1988) developed the theoretical design aspects of a controlled SDOF system with an active damper. The active damper utilized is made of prestressing tendons connected to an actuator. An optimal closed-loop control scheme using a quadratic performance index (PI) was applied to reduce the response of a structure under seismic excitation. Kobori et al. (1991a,b) presented a design method for an active mass driver (AMD) system, including the installation location, the capacity, and the stability of the system. The numerical example of an actual 10-story office building using the proposed method was shown. Chang et al. (1995) studied some design issues for buildings modeled as a SDOF system controlled by an ATMD. A closed-loop complete-feedback control algorithm was used for optimization aims. Some optimal relations for the estimation of mechanical properties of active dampers were proposed, in analogy with the classical TMDs. The ATMD design in a 10-story building frame was presented using the proposed approach. Ankireddi and Yang (1996) proposed a simplified method for the design of the ATMD based on an SDOF study of tall buildings under wind load. The procedure starts with obtaining the optimal frequency ratio and damping ratio for an equivalent TMD using analytical expressions. Subsequently, the optimal

365

366

CHAPTER 5 Design procedures for tall buildings

•

•

•

•

•

•

•

•

controller feedback gains (optimal control forces) were derived with closedform analytical expressions for the corresponding ATMD. The proposed design technique was examined for two tall buildings and analysis results show that it is accurate enough and facilitates a simple and practical design procedure. Xu (1996) introduced a method for selecting design parameters of AMDs, using a parametric study for wind-excited tall buildings. He examined the efficiency of the method with a 184-m-tall building. Yan et al. (1999) verified the analytical relations proposed by Ankireddi and Yang (1996), using some optimum algorithms for both along-wind and acrosswind excitations. They reported that the expression for the damper frequency ratio should be modified for the case of across-wind excitation. Nagashima (2001) developed an optimal displacement feedback control law, using the linear quadratic regulator (LQR), for the control of an SDOF system with an ATMD. For design purposes, the expressions of control gains were presented analytically with the help of the solution of the Riccati’s equation (that is the first-order differential equation for motion control). Moreover, analytical relations of optimal equivalent damping due to ATMDs were proposed depending on the excitation type (wind or earthquake). Ribakov et al. (2001) introduced a procedure (non-step-by-step type) for the design of building structures with active viscous damping system (AVDS) applicable to new or existing buildings. They used an active control theory (ACT) to achieve the control forces at each time step during an external excitation. The efficiency of the proposed control system was demonstrated using the numerical simulation of a 7-story building subjected to earthquakes. Cao and Li (2004) proposed control strategies for the design of ATMDs under harmonic and random forces (i.e., wind and earthquakes), leading to a simplified design method. A simple relation for the estimation of equivalent damping ratio induced by an ATMD was developed. This is very useful in the system design process. The application of the procedure to some tall building examples showed close results to those obtained from the LQR method and a better reduction on acceleration response. Du et al. (2006) developed, using a reduced-order model of a 20-story tall building, a controller design of an active mass damper to mitigate excessive vibrations induced by seismic excitation. Preumont and Seto (2008) extensively presented, in their textbook, the conceptual design of active, semiactive, and hybrid damping systems, for example, AMDs and isolated systems, using different control strategies. Various numerical examples (e.g., shear frame under seismic force) in civil structures (tall buildings) under wind and earthquake excitations could be found in the book. Ou and Li (2009) proposed a general step-by-step procedure (based on LQR control force strategy) for the optimal design of active and semiactive damping systems, for example, fluid viscous dampers, variable friction

5.4 Active, Semiactive, and Hybrid Systems

•

•

•

•

•

dampers, and MR fluid dampers. The proposed method had three general steps and it assumes identical performance for both active and semiactive control devices. Two numerical examples (20- and 76-story buildings) were analyzed. Chey et al. (2010) proposed a step-by-step design procedure of semiactivetuned mass dampers. Therein, the TMD was the entire upper portion of the building which is isolated and protected by a resettable device as the semiactive controlling system. Preumont (2011) wrote a comprehensive textbook on active control of structures. In this document, various steps for the design of active systems were briefly introduced. Xu et al. (2012) presented a detailed design process of MR dampers including the material selection, geometry design, and magnetic circuit design. The experimental tests of such dampers were performed to investigate its mechanical behavior and energy dissipation performance. Consequently, it was established that the proposed design approach is reliable for designing and optimizing MR dampers. Yang et al. (2014) proposed a simplified design method of shear-valve MR dampers considering the magnetic circuit optimization. They developed a damping force prediction model of the MR dampers and demonstrated that the simplified design procedure proposed is simple, effective, and reliable. Hazaveh et al. (2016) developed a code-based (simplified) stepwise procedures (force-based and displacement-based) for design of semiactive viscous dampers, in which the building is modeled as an SDOF system. In this procedure, the modified relations for damping reduction factor proposed by Eurocode 8 (CEN, 2004) were utilized.

5.4.2 STEP-BY-STEP PROCEDURE A simplified step-by-step design procedure, which is applicable for both active, and hybrid semiactive control of structures, is presented in this section (see Fig. 5.26). This is similar to the general procedure proposed for passive dampers shown in Fig. 5.8. Given that no code-based method is available, the main steps of the proposed procedure are based on Preumont (2011) and Ou and Li (2009). Meanwhile, some other relevant works on active control-based damping systems are reviewed within the procedure to provide a more detailed design process. Concerning the step related to the design of active control system, according to the literature, the simplest and most popular active control strategy is to use the LQR algorithm. This method needs to solve the Riccati’s equation to find the optimal gain parameters of the active (semiactive) control system. In order to solve the Riccati’s equation, users have to employ numerical software, for example, MATLAB (MathWorks Inc., 2016) or MATHEMATICA (Wolfram Research Inc., 2016). Therefore, to tackle the complexity of the use of optimization algorithms, some relations developed in the literature, for the simple (optimal) design

367

368

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.26 Step-by-step procedure for active, semiactive, and hybrid damping systems.

of active and semiactive controlling systems, are applied in the present procedure. For instance, simplified expressions are proposed by the following: •

AMD. The analytical solutions of Riccati’s equation proposed by Nagashima (2001) to obtain optimal gain parameters instead of using complicated algorithms. Moreover, additional simplified expressions were provided by Ankireddi and Yang (1996), Yan et al. (1999), and Cao and Li (2004).

5.4 Active, Semiactive, and Hybrid Systems

• •

Active system. In lieu of using the LQR approach simplified design method was proposed by Cao and Li (2004). Semiactive systems. For MR fluid dampers, simplified design expressions were proposed by Yang et al. (2014). Moreover, simplified design process for a semiactive viscous damper was introduced by Hazaveh et al. (2016) and for a semiactive-tuned mass damper (SATMD) by Chey et al. (2010).

Most of these simplified procedures are based on the construction of an equivalent SDOF. This is important because, when dealing with the MDOF systems constructing the structural matrices (M,C,K) can be difficult if the structural system is complex. Therefore, idealization of the structure as an SDOF system, based on the fundamental mode, is carried out in the proposed procedure.

5.4.2.1 Step 1: Building and site categorization In the first step, the tall building categorization (e.g., the building risk category and relative occupancy importance factor) and relative site categorization (e.g., spectral response acceleration and response spectrum) should be specified. For more details about above-mentioned information, the reader is encouraged to refer to Section 5.2.1.1.

5.4.2.2 Step 2: Select the structural system(s) Being the tall building to be designed a new structure, the proper selection of a seismic load-resisting structural system(s) is desired. To this end, Section 5.2.1.2 provides adequate details for designers based on ASCE 7-16 (ASCE, 2017a).

5.4.2.3 Step 3: Building fundamental properties and preliminary structural analyses After the definition of the main structural system, modal properties, including frequencies, fm , mode shapes, φm , generalized mass, generalized stiffness, and modal participation factors, should be estimated. For this aim, a numerical (linear) modal analysis with the use of available commercial software could help designers to compute these parameters. The required number of modal periods and mode shapes should be defined to reach at least 90% mass participation of the actual mass in each horizontal direction of the building, as recommended by ASCE 7-16 (ASCE, 2017a). It should be noted that the effective stiffness for RC structural elements should be considered in the model in order to more accurately calculate the modal parameters (see Section 5.1.3). Subsequently, the designer should analyze and design the bare structural system based on code requirements (ASCE, 2017a). From this analysis, the designer would understand if the bare structural system can resist the lateral loads without requiring excessive structural member sizes and detailing to satisfy strength, drift, and acceleration requirements. Moreover, for seismic loading inelastic behavior is expected, and this would induce damage to the structure causing reliability and economical concern as already explained in Chapter 3.

369

370

CHAPTER 5 Design procedures for tall buildings

Based on these analyses, the designer would understand the necessity to add dynamic modification system to the main structural system selected in Step 2 (Section 5.4.2.2). In the following steps, the required procedures for the design of structures will be active, semiactive, and hybrid damping technologies that will be reviewed in detail.

5.4.2.4 Step 4: Selection of performance objective In this step, the appropriate PO for the control strategy is selected. The objective may differ depending if the main is to mitigate identified deficiencies (for existing buildings) or structural response. The selection of the appropriate PO is essential for the design and evaluation of the active, semiactive, and hybrid damping system. A possible objective could be to reach a target total damping ratio and as such the optimum gain parameter(s) or control force(s) can be defined based on the analysis relations as shown in Step 8 (Section 5.4.2.8).

5.4.2.5 Step 5: Specification of the bandwidth of active and semiactive systems In this step, depending on the dominant disturbance-type exciting the structure (as determined in Step 1, Section 5.4.2.1), it is appropriate to specify the effective bandwidth for the controller system, ωb . This can be determined as the frequency range in which the amplitude of the Fourier transform of the excitation is more dominant (Preumont and Seto, 2008). For a better understanding Fig. 5.27 shows the effective excitation (harmonic) bandwidth for a typical system. It can be seen that when the excitation has a limited bandwidth, for ω , ωb , the contribution of higher order frequency modes in the overall response, and consequently in the controller system, becomes insignificant. Therefore, the knowledge about the effective bandwidth can help the designer in choosing which vibration modes to control; this is useful when dealing with model reduction in the full structure.

5.4.2.6 Step 6: Damper type, configuration, and distribution of sensors and actuators 5.4.2.6.1 Device type The choice of type of active control systems depends upon several factors, for example, the PO, excitation type, and structural configuration. As discussed in detail in Section 4.3 (Chapter 4) the designer can choose between several different devices, for example, AMD, AVDS, and semiactive base-isolated systems. The designer choice is based on the most suitable system to achieve the desired structural performance. One of the major criteria for the selection of active-based controller devices is the type and level of excitation, for example, wind excitation and small/largescale earthquakes. Preumont and Seto (2008) reported that AMD and hybrid mass damper (HMD) are useful when dealing with wind excitation or small scale of earthquake. Using active- and hybrid-controlled bridge between two buildings is

5.4 Active, Semiactive, and Hybrid Systems

FIGURE 5.27 (A) Fourier spectrum of the excitation F with a limited frequency content ω , ωb . (B) Dynamic amplification of mode m such that ωm , ωb and ωk cωb . Adapted from Preumont, A., Seto, K., 2008. Active Control of Structures. John Wiley & Sons. Ltd, Chichester, UK.

effective, when the structure is excited by small/large-scale earthquakes. Also, semiactive brace-damper and semiactive base-isolated systems (e.g., adjustable oil damper and MR damper) can be reliable under large-scale earthquakes.

5.4.2.6.2 Device location Since active/semiactive control system contains sensors (for monitoring the structural response) and actuators (for controlling the structural response), their location is the most important factor influencing the performance of the control system. The actuators are suggested to be placed where their ability to control the modes is the largest (Preumont, 2011). A control system where actuator and sensor are attached to the same DOF is called a collocated control system (Preumont and Seto, 2008; Preumont, 2011). In this case, actuator and sensor should also be dual, that is, the actuator force must be associated with a translation sensor (e.g., measuring displacement, velocity, or

371

372

CHAPTER 5 Design procedures for tall buildings

acceleration). Therefore, the product of the actuator signal and the sensor signal represents the energy (power) exchanged between the structure and the control system. The location of actuators may also depend upon the type of control system to be designed. For instance, single AMD which is usually installed at the top of buildings requires actuators/sensors at the same level (Abdullah et al., 2001). Also, active tendon mechanisms can be easily installed at discrete locations on civil engineering structures (Abdullah et al., 2001). The optimum placement of sensors and control systems, such that their efficiency can be maximized, was studied by Brown et al. (1999) for shear frames subjected to earthquake and wind excitations. Using a multiobjective linear quadratic Gaussian methodology based on Pareto optimal controllers, they concluded that: • • •

The use of many actuators is preferable. The actuator placement is much more important than sensor placement. The number and location of actuators depend on the properties and configurations of the structure.

Abdullah et al. (2001) suggested the simultaneous placement of sensor/actuators to obtain a more economical and desirable design. For an accurate placement of sensors and actuators in tall buildings, there are various methods such as GAs (Liu et al., 2003, Tan et al., 2005; Cha et al., 2012) and artificial neural networks (Amini and Tavassoli, 2005). It is worth noting that the review of these more complicated techniques is beyond the scope of this book. Therefore, in the following step, a simplified model is described.

5.4.2.7 Step 7: Build simplified and reduced model To build the simplified model, the first phase is to compute the modal properties (i.e., natural frequencies (periods) and mode shapes) through a linear modal analysis as done in Step 3 (Section 5.4.2.3). If the actuators are considered in the frequency band of interest, their effects can be neglected temporarily in the model and taken into consideration on the system performance control after the design is completed. Otherwise, they should be considered in the model before the design of controller (Preumont, 2011).

5.4.2.7.1 Model reduction In general, structures are distributed systems by nature, and design of a controller system may be complicated for such structures. In contrary, a reduced-order model based on lumped parameter systems is more suitable for the design of a controlling system in a physical state space, since the number of modes is reduced (Seto et al., 1998). One of the major issues of utilizing such an approach is the elimination of the spillover phenomenon, that is, the interference of higher modes on the controlled ones (Balas, 1978). This can be achieved by allocating actuators or sensors in correspondence with neglected higher mode nodes (further discussion will be given in the following sections).

5.4 Active, Semiactive, and Hybrid Systems

To obtain a reduced-order model the first stage is to determine the number and positions of modeling points (i.e., locations of lumped masses). The choice of these points is based on the strategy utilized in such a way that the main structure is observable (using sensors) and controllable (using actuators), to have a meaningful controller design (Preumont and Seto, 2008). Assuming that a tall building may be seen as a continuum (beam-like) cantilever system, the modeling points can be determined using the first four mode shapes of the structure (Fig. 5.28 (Preumont and Seto, 2008)). In Fig. 5.28, φqi and φri are the ith mode shape components at the location of control (actuator-induced) force (i.e., point q) and sensor (i.e., point r), respectively. If control force and sensor points are zeros, the system is uncontrollable and unobservable; while if not zeros, the system is controllable and observable. Accordingly, if the control force (actuator) is applied to a modal node (i.e., where the mode shape value is zero, φ 5 0) than this mode is uncontrollable. Moreover, if a sensor is mounted to a modal node that mode is then unobservable (Preumont and Seto, 2008). As it can be observed from Fig. 5.28, the first and second modes are both controllable and observable since the control force and sensor are located at nonzero nodes of the corresponding modes. In contrary, the third mode is observable but uncontrollable, because the control force (point q) is placed at a zero value of the mode while the sensor (point r) is not. Inversely, the fourth mode is controllable but unobservable due to the location of the actuator at nonzero node and the sensor at zero node of the fourth mode. Based on these considerations, a general method to reduce the full model to a simplified one can be summarized here as follows: • • • •

Analysis and determination of the mode shapes of the reference structure that will be controlled (e.g., first two modes) Selection of the zero nodes of the lowest order mode that is not controlled (e.g., third mode) Placement of the lumped masses at the select nodes (at zero nodes of the lowest order mode), resulting in an n-DOF system Determination of the mass and stiffness constants through sensitivity analysis

Thus, it can be concluded that if the objective is to just control the first two modes, two modeling points (i.e., the points that are both controllable and observable) are required. These points can be located at the zero nodes of the third mode, and in this way the actuator and sensor can be placed confidently. Note that in doing this the third mode becomes uncontrollable and unobservable.

5.4.2.7.2 Modeling the SDOF system To control tower-like structures, the behavior can be simplified with an SDOF model (Preumont and Seto, 2008). In this case, design of active dynamic absorbers, such as AMDs, becomes easier. For constructing an SDOF system, controlling the first two modes, it is necessary to understand the relationship between modeling points and the corresponding SDOF systems. There are two possible modeling

373

FIGURE 5.28 Relationship between controllability/observability and locations of sensors and actuators. Adapted from Preumont, A., Seto, K., 2008. Active Control of Structures. John Wiley & Sons. Ltd, Chichester, UK.

5.4 Active, Semiactive, and Hybrid Systems

(A)

Modeling points

(B)

u1 m1 u2 m2 k1 k2

FIGURE 5.29 SDOF model for tower-like structures. Adapted from Preumont, A., Seto, K., 2008. Active Control of Structures. John Wiley & Sons. Ltd, Chichester, UK.

approaches as shown in Fig. 5.29. In Fig. 5.29A, the modeling point is at the top. In this case, locating the actuator (control force) and sensor at this point leads to a controllable and observable controlling system for the first mode. Moreover, this modeling point is not coincident with the zero node of the second mode leading to effectively control it. Instead in Fig. 5.29B, the selected modeling point is located at the zero node of the second mode. Hence, application of the control force and sensor at this point does not let to control the second mode, but it is still useful in controlling the first mode. Therefore, the choice of the modeling point (position of the lumped mass) depends upon the designer’s need to control either modes or only the first mode. In Fig. 5.29, m1 , m2 , and k1 , k2 are the masses and stiffnesses estimated at two types of modeling points, respectively. These values can be easily achieved using the mass response method (Seto et al., 1987). It is useful to know that if the modeling point is selected at the top structure (Fig. 5.29A), smaller mass and stiffness are obtained (Preumont and Seto, 2008). Hence, the best mounting location of the AMD or HMD controller is at the top of the structure for controlling the first vibration mode (Preumont and Seto, 2008).

5.4.2.7.3 Higher order DOF systems In addition to the SDOF system, it is also possible to reduce the full structure as an equivalent system with two, three, and four lumped masses: •

Two-DOF system (Seto et al., 1995) is useful to consider the behavior of the first two bending modes of the structure.

375

376

CHAPTER 5 Design procedures for tall buildings

• •

Three-DOF model (Kar et al., 2000) is applied for a broad building structure, accounting for twisting modes in addition to bending ones. Four-DOF model (Preumont and Seto, 2008) can be used for bridge tower structures to control the first four vibration modes, two bending and two twisting ones.

5.4.2.7.4 Spillover phenomenon In ACT, when dealing with a reduced model of a structure, a problem called spillover may occur. This problem is due to the interference between the controlled modes and uncontrolled ones (higher modes). In other words, when applying the control force into the controlled modes, the actuator may also excite the higher modes that are not included in the reduced model. Similarly, sensors can have spillover problems since they may not only capture the information related to the controlled modes, but also that of the uncontrolled modes. The spillover may be a dangerous phenomenon, leading to local damages. Therefore, avoiding or suppressing spillover is a very important task for a successful control. There are different ways to reduce/avoid spillover problem, such as: •

•

•

•

Modal filtering. It is essentially based on the separation of the information for the controlled modes from uncontrolled ones. This approach needs the same number of sensors and actuators as that of the controlled modes (Balas, 1982). Direct feedback. It requires the placement of sensors and actuators at the same locations (i.e., collocation). Significant control effects could not be anticipated with this method, but the stability of this control system is guaranteed (Balas, 1979). Low-pass filtering. It is based on cutting off either the captured sensor signals or the control (actuator) forces to reduce the effects of higher uncontrolled modes. This method is simple but has often failed in controlling multiple modes (Hori and Seto, 2000). Robust control. Filtered LQR and H-infinity (based on minimizing two weighting functions (Preumount and Seto, 2008)) state feedback controls are some robust control methods used to mitigate the effects of uncontrolled higher modes (Hori and Seto, 2000; Preumount and Seto, 2008).

It is worth noticing that, in the following procedure, the SDOF reduced modeling approach is selected. This will simplify the design of the controlling system as shown in the subsequent step.

5.4.2.8 Step 8: Design of the controller system In this step, several simplified design approaches depending on the type of controller system (e.g., ATMDs, MR dampers, and semiactive dampers) are presented. For ATMDs, which are most commonly used by active controlling devices in tall buildings, three approaches are represented. Moreover, a simplified approach for design of MR dampers, which are widely used as semiactive control devices, is provided. Lastly, simplified design guidelines for semiactive viscous

5.4 Active, Semiactive, and Hybrid Systems

dampers and TMDs are addressed. All the approaches contain several substeps addressing simplified relations for which the design parameters can be determined.

5.4.2.8.1 Design of ATMD For ATMD system, there are several works that address simplified design relations to obtain optimal control gains without dealing with complex optimization algorithms. Three simplified approaches are reviewed in this section: Approach 1 (based on Ankireddi and Yang (1996) and Yan et al. (1999)); Approach 2 (based on Nagashima (2001)); Approach 3 (based on Cao and Li (2004)). Approach 1 (Ankireddi and Yang, 1996). In this approach, the building is modeled as an SDOF system, dominated by the first-mode response. The approach is composed of two stages in which, first, the optimal characteristics of a passive TMD is predetermined; then, the active control system is designed accordingly. This method is applicable for buildings under wind loads and is developed based on the complete-feedback (namely, feedback of displacement, velocity, and acceleration) control of the system. Note that the main structure is assumed undamped (i.e., negligible inherent damping). The sequence of parameters to define is the following: •

•

Selection of a mass ratio μ In this step, a trial mass ratio, μ, as the ratio between the auxiliary mass and main system mass (generalized mass associated with the suppressed mode of the building) can be selected. For a proper selection of this parameter, one can refer to Step 7 of the mass damping procedure (Section 5.2.2.8.1). Specification of optimal frequency ratio fopt and supplemental damping ratio ζ d;opt In this stage, using the selected mass ratio and assuming a negligible inherent damping, the following relations, proposed by Ankireddi and Yang (1996), can be utilized to determine the ATMD optimal values of frequency and damping ratio (for a deeper discussion refer to Step 7, Section 5.2.2.8.2). For buildings under along-wind excitation: sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 21μ μð4 1 3μÞ ;ζ d;opt 5 fopt 5 8ð1 1 μÞð2 1 μÞ 2ð11μÞ2

(5.182)

These relations are also verified by Yan et al. (1999). For buildings under across-wind excitation, the following expression is suggested for the optimal frequency by Yan et al. (1999) to achieve more accurate results:   fopt 5 134:6f12 2 87:74f1 1 11:9 μ2   2 26:82f12 2 17:71f1 1 3:359 μ   2 0:029f12 2 0:019f1 1 0:996

Here, f1 is the fundamental frequency of building.

(5.183)

377

378

CHAPTER 5 Design procedures for tall buildings

•

•

Determination of mechanical properties of auxiliary mass Having determined the optimal frequency ratio and damping ratio (Eqs. 5.182 and 5.183), the stiffness, kd , and damping constants, cd , of the auxiliary mass should be specified as: kd 5 md ω2d

(5.184)

cd 5 2ζ d;opt md ωd

(5.185)

where md and ωd are generalized mass and frequency of the auxiliary mass, respectively. After determination of optimal properties of a passive mass damper, the next phase of the design should be based on calculating active controlling features. Selection of acceleration feedback gain ratio μg In this step, a parameter, called acceleration feedback gain ratio, μg , should be selected. This parameter controls the ratio between the gain of acceleration feedback, mg , adopted for the active control and the generalized main mass, m1 : μg 5

mg m1

(5.186)

A selection criteria (effective range), such that the ATMD system function is stable, was proposed by Ankireddi and Yang (1996) as: 2

•

μd , μg # 0 11μd

which means that μg has a negative value. Calculation of the optimal controller velocity and displacement feedback gain coefficients εg;opt and ψg;opt The optimal values of velocity feedback gain, εg;opt , and displacement feedback, ψg;opt , as proposed by Ankireddi and Yang (1996) and Yan et al. (1999) can be calculated as follows: cd 1 cg εg;opt 5 5 m1 ω 1 ψg;opt 5

•

(5.187)

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi μg μ2 ð413μÞμ3 1 2 ð11μÞ 4ð11μÞ3

μ2 1 2μ 1 2μg 1 2μg μ kd 1 kg 5 2 m1 ω1 2ð11μÞ2

(5.188)

(5.189)

Determination of optimal gains of velocity and displacement feedback Having determined the velocity and displacement feedback coefficients in the previous step, the optimal control gains can be derived from the following set of equations:   cg 5 εgopt 2 2ζ s;opt fopt μ m1 ω1

2 μ m1 ω21 kg 5 ψgopt 2 fopt

(5.190) (5.191)

5.4 Active, Semiactive, and Hybrid Systems

•

Determination of optimum control force An optimum active control force, uopt , is desired, because an increase in this force does not necessarily mean an increase in ATMD effectiveness: uopt 5 2mg x€ d 2cg x_ d 2kg xd

(5.192)

where xd is the relative displacement of the active mass damper. Approach 2 (Nagashima, 2001). For a structure controlled by an ATMD system, the active control law (LQR method) can be simplified to be composed of two different feedback gains: one for the structure displacement and one for the auxiliary mass velocity (Nagashima, 2001). This is considered as “optimal displacement feedback control” that may be useful for design of ATMD systems. Hence, using the simplified expressions developed by Nagashima (2001), the damping design parameters of the ATMD system can be proposed in closed-form solution for an undamped SDOF structure controlled with an ATMD. The sequence of parameters to define is the following: • •

•

•

Selection of a mass ratio μ The mass ratio is selected in the same way as shown in Approach 1. Selection of a target supplemental damping ratio A target supplemental damping, ζ d , can be selected based on the recommendations given for the step-by-step procedure for mass damping systems (Step 5, Section 5.2.2.6). It is worth mentioning that a desirable damping ratio may be more easily specified using the analytical expressions of the optimal damping ratio for TMDs in undamped SDOF systems (refer to Table A.1 (Appendix A)). Calculation of auxiliary mass md Based on the mass ratio, μ, selected, the auxiliary mass, md , can be computed similarly to what was shown in Eq. (5.112). Estimation of the proposed displacement feedback gain (G1 ) In this step, based on the simplified relations developed by Nagashima (2001), an active control parameter called displacement feedback gain (G1 ) can be determined for both wind and earthquake excitations:
Wind excitation  4ζ 2 G1 5 1 2 d md ω21 μ




(5.193)

where ω1 is the frequency of fundamental mode, and 0 # G1 # md ω21 . Note that for G1 . md ω21 the added damping ratio expected by the active system is negligible, that is, ζ d 5 0. Earthquake excitation  G1 5 m ω2

1 4ζ 2 2 d md ω21 μ 11μ m ω2

(5.194)

d 1 d 1 where 0 # G1 # 1 1 μ. But, for G1 . 1 1 μ the damping ratio expected by the active system is negligible, that is, ζ d 5 0.

379

380

CHAPTER 5 Design procedures for tall buildings

•

Calculation of the proposed velocity feedback gain (G4 ) of auxiliary mass stroke Similar to the control gain determined in the previous step, Nagashima (2001) proposed an active control parameter called velocity feedback gain (G4 ) that can be estimated as follows: sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  2 ω2 G4 5 2ζ d ωd md 2 2μ d2 G1 1 2ζ d ωd md ω1

•

where G1 $ 0 and ωd is the frequency of the auxiliary mass. Feedback control low Having calculated the control gain, it is possible to compute the feedback control law, u, as a function of time as follows: uðtÞ 5 G4 x_ d 1 G2 x1

•

(5.195)

(5.196)

Here, x_ d and x1 are, respectively, the velocity of auxiliary mass stroke and displacement of the main mass (SDOF) system. Determination of frequency of auxiliary mass ωd According to Nagashima (2001), two PIs are used for the optimization of feedback gains: displacement performance index (DPI) and velocity performance index (VPI). Therefore, the optimal frequency, ωd , of the ATMD for both of these indices can be determined as follows:
Optimization based on DPI ωd 5

ω1 11μ

(5.197)

where this tuning frequency coincides with the well-known optimal tuning condition for a passive TMD developed by, for example, Den Hartog (1956), Warburton (1982a,b), and Chritopulos and Filiatrault (2006) (refer to Chapter 4, Section 4.1.2.1).
Optimization based on VPI ω1 ωd 5 pffiffiffiffiffiffiffiffiffiffiffi 11μ

•

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi μ G4 11 2 md ω21

(5.198)

Determination of mechanical properties of auxiliary mass Having determined the amount of auxiliary mass (md ) and frequency (ωd ), the stiffness and damping constants of the ATMD systems are given, respectively, in Eqs. (5.184) and (5.185). Nagashima (2001) emphasized that the optimization based on the DPI is practical and the ATMD performance is improved as the displacement feedback gain G4 increases. However, the system performance using the VPIbased optimization is limited (increase in the stiffness of the auxiliary mass with G4 and not improving control performance even for large G4 ), leading to an inappropriate control strategy.

5.4 Active, Semiactive, and Hybrid Systems

Approach 3 (Cao and Li, 2004). Cao and Li (2004) developed a simplified approach for the design of control strategy, based on LQR technique, which accounts for the inherent damping of the main system. The sequence of parameters to define is the following: • •

Selection of a mass ratio μ The mass ratio is selected in the same way as shown in Approach 1. Determination of optimal frequency ratio and damping ratio In order to be able to determine the optimal gains for control feedbacks, based on the mass ratio and inherent damping, the optimal values of frequency ratio, fd;opt , and supplemental damping ratio, ζ d;opt , can be calculated using expressions, described in Table A.2 (Appendix A), depending upon the type and position of the external excitation applied on the primary (SDOF) system. Alternatively, Li et al. (1999) proposed to utilize the following simplified relationship: pffiffiffi ζ d;opt 5 0:5 μ

•

Calculation of expected frequency and added damping of ATMD Given the optimal frequency ratio and added damping ratio from the previous step, the frequency, ωd , and added damping, ζ d , due to the passive damping system can be simply calculated as: ωd 5fd;opt ω1 ;

•

(5.199)

ζ d 5 ζ d;opt

(5.200)

Note that, alternatively, the damping factor can be selected based on a certain level of reduction in a desired response. Determination of optimal control gains The expressions developed for the optimal gains usually depend on the excitation type. Here, the relations for harmonic excitations (with frequency ω and angle phase θ) and random forces (white-noise) are presented.
Harmonic excitations (xðtÞ 5 Xsinωt) (Cao and Li, 2004) 1. Displacement feedback (G3 5 G4 5 G5 5 G6 5 0; G1 ; G2 6¼ 0)   μ 2ζ s ωd ωB 2 ω 2 Asinðθ2 2 θ1 Þ ω21

(5.201)

ω2 ω 1 2ζ s cosðθ2 2 θ1 Þ 2 1 ωd ω21

(5.202)

G1 5 G2 5

2. Velocity feedback (G1 5 G2 5 G5 5 G6 5 0; G3 ; G4 6¼ 0)   ω2 Acosðθ2 2 θ1 Þ 1 ω2 2 ω2d B 2ζ 1 ω1 ωAsinðθ1 2 θ2 Þ  2  2 ω A 1 ω 2 ω2d Bcosðθ2 2 θ1 Þ 21 G4 5 2ζ s ωd ωBsinðθ1 2 θ2 Þ G3 5 μ

3. Acceleration feedback (G1 5 G2 5 G3 5 G4 5 0; G5 ; G6 6¼ 0)

(5.203) (5.204)

381

382

CHAPTER 5 Design procedures for tall buildings   2ζ s ωd ωB G5 5 μ 21 Aωsinðθ2 2 θ1 Þ G6 5

where

ω2d ωd 2 2ζ s cosðθ2 2 θ1 Þ 2 1 ω2 ω

Λ21 v20 1 μ2 ω4 B2 2 A02 θ2 5 cos 2 2Λ1 v0 μω2 B  2 μω B 2ζ 1 ω1 ω sinθ2 1 tan21 2 θ1 5 sin21 A0 ω1 2 ð1 1 μÞω2 21

with

(5.205) (5.206)



qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  2  2 0 A 5 A ω21 2 ð11μÞω2 1 2ζ 1 ω1 ω

(5.207) (5.208)

(5.209)

Here, v0 is the harmonic force amplitude; A and B are positive (static) displacement amplitudes of main structure (SDOF) and ATMD system, respectively. Cao and Li (2004) suggested to use arbitrary values for parameter A (e.g., A 5 0.8, 1, 1.2) and determine the parameter B by satisfying the following limiting conditions:   Λ1 v0 2 μω2 B # A

Pn

(5.210) (5.211)

Si φ2i1

where Λ1 5 , in which i denotes the number of story; n is the m1 number of building floors; Si is the static amplitude of wind load at ith floor; φi1 is the modal value associated with the first mode and ith floor. Random excitation (Cao and Li, 2004) When the structure is excited by a random force, for example, whitenoise, the optimal control gains (G1 to G6 ) can be extracted from the optimization of displacement responses variant by solving the following set of linear equations (for further explanation about this set of equations interested, readers should refer to the publication by Cao and Li (2004)): i51




  Λ1 v0 2 A0  # μω2 B

8 @A4 @B4 > > B4 2 A4 50 > > > @G @G 1 1 > > > > > @A4 @B4 > > 2 A4 50 > B4 > > @G @G 2 2 > > > > > @A4 @B4 > > 2 A4 50 B > > < 4 @G3 @G3 @A4 @B4 > > B4 2 A4 50 > > > @G @G 4 4 > > > > > @A4 @B4 > > B 2 A4 50 > > > 4 @G5 @G5 > > > > > @A4 @B4 > > B 2 A4 50 > > : 4 @G6 @G6

(5.212)

5.4 Active, Semiactive, and Hybrid Systems

The fundamental parameters required for the derivation of optimal gains are:   A4 5 a0 2 a4 a21 1 a1 a2 a3 2 a0 a23

(5.213)

B4 5 a4 ½b0 ða3 a2 2 a4 a1 Þ 1 a0 ða3 b1 1 a1 b2 Þ

(5.214)

where

8 > > > >
> > a3 5 2ð1 2 G3 1 G6 Þζ 1 ω1 1 2ð1 1 μÞð1 1 G4 Þζ s ωd > : a4 5 1 2 G5 1 ð1 1 μÞG6 8 > b0 5 Λ1 2 ð11G2 Þ2 ω4d > < 2 b1 5 2 2Λ1 ð1 1 G2 Þð1 1 G6 Þω2d 1 4Λ1 2 ð11G4 Þ2 ω2d ζ s 2 > b2 5 Λ1 2 ð11G6 Þ2 > : b3 5 0

•

(5.215)

(5.216)

Determination of supplemental damping ratio due to the active controller The supplemental damping ratio due to the active control system can be estimated using the simplified relation proposed by Cao and Li (2004) as follows: ζd 5

a4 Λ21 A4 2ω31 B4

(5.217)

The active control added damping ratio, ζ d , shall be checked against the added damping ratio, ζ d , as follows: ζ d $ ζd

•

(5.218)

If the above inequality is not satisfied or ζ d is significantly larger than ζ d , the designer can change the controlling parameters A4 or B4 by changing the mass ratio, μ, supplemental damping ratio, ζ d , or frequency, ωd , of the ATMD system. Specification of mass, stiffness, and damping of ATMD Having obtained the mass ratio (μd ) and frequency (ωd ), the mass, stiffness, and damping constants of the ATMD system are calculated as shown in Eqs. (5.112), (5.184), and (5.185), respectively.

5.4.2.8.2 Design of semiactive systems Concerning semiactive controlling systems, Ou and Li (2009) reported that various semiactive dampers (e.g., MR dampers and variable friction dampers) can be designed in the same manner as the active dampers of the same type. This assumes that the response of a building controlled with active control systems can be considered to be the same as that with semiactive devices. First the design strategy, proposed by Ou and Li (2009), considers to determine active control

383

384

CHAPTER 5 Design procedures for tall buildings

forces using the LQR algorithm, according to the expected reduction in a response; then, replaces the active devices by semiactive ones by taking the following assumptions: • •

The maximum output force of the semiactive device is identical to that of the active control systems. The response of the structure with the semiactive systems is the same as that with active devices.

As mentioned earlier, for the determination of active control forces, it is required to employ complicated algorithms (e.g., LQR), and this results in a timeconsuming and complex design. Therefore, more simplified design procedures are presented in this section for the most practical semiactive systems: MR fluid dampers, viscous dampers, and TMDs. Design of MR dampers. In general, the working mode of MR dampers can be categorized as follows: shear, valve, extrusion, and shear-valve modes. Among these, the shear-valve mode has been extensively employed in the design of MR dampers for civil engineering structures because of its efficiency and geometry simplicity (Yang et al., 2014). The MR damper design process generally consists of the following phases: material selection, geometry, and magnetic circuit design. All these aspects are addressed in the following based on the recommendations of Yang et al. (2014). •

•

Material selection Suitable materials for MR dampers are generally based on two objectives (Xu et al., 2012):
Selection of a type of MR fluid with low apparent viscosity and proper magnetic saturation yield strength
Selection of cylinder and piston materials in such a way that the saturation induction density is higher than the magnetic field intensity when MR fluid reaches magnetic saturation yield strength As an example, fluid MRFXZD08-01 which has antisettlement properties and low apparent viscosity in low-frequency vibration has been utilized in the past (Xu et al., 2012). Moreover, the DT4 electrical pure iron and No. 45 steel have been adopted for manufacturing the piston and the cylinder, respectively (Xu et al., 2012). Another practical MR device utilizes carbon steel for piston and cylinder and MRF-J01 fluid (Yang et al., 2014). Geometric and properties selection/design The main geometric parameters that define MR dampers can be computed as follows (Fig. 5.30):
Total effective length of the damper is expressed as follows (Yang et al., 2014): L5

2nr 2 B1 pffiffiffiqffiffiffiffiffiffiffih 2 2rB 1 3 2B μ I Bd 0

(5.219)

FIGURE 5.30 Schematic diagram of a shear-valve MR damper. Adapted from Yang, D., Lu, Z., Zhu, H., Li, Z., 2014. Simplified design method for shear-valve magnetorheological dampers. Earthquake Eng. Eng. Vib. 13, 637
652.

386

CHAPTER 5 Design procedures for tall buildings

where n is the number of piston sections; B1 is the saturation value of magnetic induction of the magnetic core (1.5 for carbon steel; Yang et al. (2014)); B2 is the saturation value of magnetic induction of MR fluids (1 for fluid MRFXZD08-01 (Xu e al., 2012) and 0.6 for fluid MRF-J01 (Yang et al., 2014)); I is the maximum value of current input; μ0 is the permeability of air; d is the diameter of electromagnetic coils; h is the width of damping path; and r is the radius of magnetic core. Note that the resulting L is a function of h and r at this stage (i.e., h and r are unknown at this level).
Average circumference of the damping path is estimated as follows (Yang et al., 2014): sffiffiffiffiffiffiffiffiffiffi pffiffiffi 2B2 h D 5 2r 1 h 1 3 d μ0 I 0

(5.220)

0




Here, D is a function of h and r. Effective cross-sectional area of the piston is estimated as follows (Yang et al., 2014): !2 pffiffiffi sffiffiffiffiffiffiffiffiffiffi 3 2B2 h Ap 5 π d1r 2 πr02 μ0 I 2

(5.221)

where Ap is a function of h and r. Moreover, r0 is the radius of piston plate, which is estimated as follows: r0 $




rffiffiffiffiffiffi F πσ

where σ is the yielding stress of material of the piston shafts. Depth of groove accommodating the electromagnetic coil is estimated as follows (Yang et al., 2014): pffiffiffi sffiffiffiffiffiffiffiffiffiffi 3 2B2 h h1 5 d 2 μ0 I




(5.223)

Damping path effective length, L1, and width of the groove, L2, accommodating the electromagnetic coils are estimated is as follows (Yang et al., 2014): L1 5

r 2 B1 2ðr 1 h1 ÞB2

L2 5 2h1




(5.222)

(5.224) (5.225)

Thickness of the cylindrical housing estimated is as follows (Yang et al., 2014): t5

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi B2 2ðr 1 h1 ÞL1 1 ðr1h1h1 Þ2 2 r 2 h 2 h1 B2

(5.226)

5.4 Active, Semiactive, and Hybrid Systems

To control the performance requirements of the MR damper, two additional parameters need to be defined: dynamic range, β, and controllable force, F. The dynamic range is the adjustment factor of damping force which equals to the ratio of its maximum to its minimum (that should be as large as possible (Xu et al., 2012; Kasemi et al., 2011)). The controllable force usually consists of two components: valve and shear (Yang et al., 2014). The appropriate choice of these parameters is subject to iteration since they are functions of several parameters, which are as follows: 8 12ηLA2p 3Lτ y > > > Ap F5 u1 > < πD0 h3 h > > > > :

•

β5

πD0 τ y h2 4ηAp u

(5.227)

where τ y and η are the shear yield stress and viscosity of MR fluid, respectively; and u_ is the relative velocity between piston and cylindrical housing (Xu et al. (2012) proposed the velocity of 100 mm/s and Yang et al. (2014) used 80 mm/s). At the beginning an initial guess should be tried, and in literature several recommendations can be found, such as:
Yan et al. (2014) suggest that β values range between 7.7 and 22 are based on experimental testing for civil applications.
Yang et al. (2000, 2001, 2002a,b) and Fujitani et al. (2002) designed and fabricated a full-scale MR damper with a maximum damping force of F 5 200 kN.
Xu et al. (2012) developed the design of an MR damper system for mitigation of earthquake in civil engineering for F 5 200 kN and β . 15 Magnetic circuit design The objective of magnetic circuit design is to determine the number of electromagnetic coils, N, so that the magnetic induction density in the damping path, generated by magnetic circuit, is more than the magnetic field intensity (Xu et al., 2012). This number can be estimated by (Yang et al., 2014): N5

2B2 h μ0 I

(5.228)

It can be seen that N is only dependent on h and independent on the other geometric parameters. Design of semiactive viscous dampers. In this section, the design of a semiactive viscous damper (Fig. 5.31) is presented as proposed by Hazaveh et al. (2016). The orifices existing in the device may be opened or closed depending on the direction of velocity and displacement at each time step (sensors are installed in the structure to record these responses). Correspondingly, the minimum and maximum damping can be achieved when the orifices are closed or opened, respectively.

387

388

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.31 Schematic representation of semiactive devices attached to an SDOF system. Adapted from Rodgers, G.W., Mander, J.B., Chase, J.G., Mulligan, K.J., Deam, B.L., Carr, A., 2007. Reshaping hysteretic behaviour—spectral analysis and design equations for semi-active structures. Earthquake Eng. Struct. Dyn. 36 (1), 77
100.

FIGURE 5.32 Schematic hysteresis for (A) 1-4 device, (B) 1-3 device, and (C) 2-4 device. Adapted from Hazaveh, N.K., Pampanin, S., Rodgers, G.W., Chase, J.G., 2014. Novel semi-active viscous damping device for reshaping structural response. In: Conference: 6WCSCM (Sixth World Conference of the International Association for Structural Control and Monitoring).

Based on such a device, three different device control laws (called 1-4 device, 1-3 device, and 2-4 device) can be presented (see the corresponding hysteretic loops in Fig. 5.32 (Hazaveh et al., 2014)). The 1-4 law provides damping in all four quadrants recreating a response similar to passive viscous dampers (Fig. 5.32A). The 1-3 law provides damping only in the first and third quadrants, resisting away from

5.4 Active, Semiactive, and Hybrid Systems

equilibrium (Fig. 5.32B). The 2-4 law provides damping only in the second and fourth quadrants, resisting only toward equilibrium (Fig. 5.32C). The simplified design herein is mainly based on the 2-4 device (Fig. 5.32C) which is simultaneously effective in reducing displacement and base-shear demands. Note that the main structure is idealized as an SDOF system. •

•

Selection of supplemental damping ζ d The supplemental damping of semiactive device, ζ d , can be selected. The effective damping of the device 2-4 that it is added to structure is different from the nominal damping capacity of the device, determined as shown in the next step. Estimation of nominal damping ζ dn The nominal supplemental damping for 2-4 devices can be determined given through one of the following methods (Hazaveh et al., 2016) (based on interpretation of numerical results):
Linear approximation method 8 >
 pffiffiffiffiffi : ζ dn 5 0:07 2 0:02 Bζ Bζ ð24:54Þ for 2:7 , T1 # 5 s




where Bζ is the damping reduction factor (as discussed in Section 5.2.1.4.1) and T1 is the SDOF fundamental period. Eurocode 8 (EC8, 2004) as modified by Priestley et al. (2005, 2007) The following relation (Hazaveh et al., 2016) is proposed to straightforwardly estimate the total nominal damping of devices based on the effective damping: ζ dn 5 4:93ζ d 2 0:078

•

(5.229)

(5.230)

Note that if the ζ dn resulted from two above equations are different, an average of two values can be considered as the design damping (Hazaveh et al., 2016). Determination of damping coefficient C From Fig. 5.33 it can be seen that the hysteretic loop of the 2-4 semiactive viscous dampers (Fig. 5.33A) is half of the area associated with one to four

FIGURE 5.33 Hysteretic loops of (A) 2-4 device and (B) 1-4 device. Adapted from Hazaveh, N.K., Rodgers, G.W., Pampanin, S., Chase, J.G., 2016. Damping reduction factors and code-based design equation for structures using semi-active viscous dampers. Earthquake Eng. Struct. Dyn.

389

390

CHAPTER 5 Design procedures for tall buildings

device (Fig. 5.33B). Accordingly, the relationship between supplemental equivalent damping ratios is determined as follows (based on Eq. (3.60) (Chapter 3)):  1 1 Ed ζ d;224 5 ζ d;124 5 2 2 4πEs

(5.231)

Assuming a harmonic motion (Ed 5 πcωU02 and Es 5 k0 U02 =2, where k0 is the structural stiffness of the SDOF system (main structure) and U0 is the amplitude of harmonic motion) and considering the same damping constant for both systems (1-4 and 2-4 device), after some manipulations, the following expression can be obtained: ζ d;224 5

 1 Tc 2 4πm

(5.232)

Here, T and m are, respectively, the natural period and mass of SDOF system. Consequently, considering ζ d;224 5 ζ d , the equivalent damping constant is given by: C5

•

8πm ζ T d;224

(5.233)

Determination of nominal damping coefficient The choice of device configuration and distribution is basically similar to the case of passive (viscous) dampers (Section 5.2.1.6). Considering a uniform distribution of devices along the building height the nominal damping coefficient associated with a single damper is simply calculated by: c5

c N

(5.234)

Therefore, based on the performance function of 2-4 semiactive viscous dampers, the damping force in such devices is expressed by:  FD 5

•

cu_ d 0

for sgnðud Þ 6¼ sgnðu_ d Þ for sgnðud Þ 5 sgnðu_ d Þ

(5.235)

Determination of physical dimensions The fundamental design parameters of such a semiactive system are piston diameter, D, individual chamber length, L0 , and maximum piston displacement, δ, as shown in Fig. 5.34. Mulligan (2007) states that the piston displacement should be less than chamber length (i.e., δ # L0 Þ. The use of these parameters is helpful in controlling the stiffness of the device, Kd . The verification of such parameters should be conducted using experimental tests as shown in Step 9, in Section 5.4.2.9. The piston diameter can be approximated using the expression proposed by Mulligan (2007), as follows: sffiffiffiffiffiffiffiffiffiffiffiffiffi 2L0 Kd D5 πγP0

(5.236)

5.4 Active, Semiactive, and Hybrid Systems

FIGURE 5.34 Design parameters of 2-4 semiactive damper. Adapted from Mulligan, K.J., 2007. Experimental and analytical studies of semi-active and passive structural control of buildings. Ph.D. Thesis, Department of Mechanical Engineering, University of Canterbury, Christchurch, New Zealand.

where γ is the ratio of specific heats (1.4 for air) and P0 is the initial chamber pressure (typically atmospheric pressure). The chamber length, L0 , is typically constrained by the application and it is determined by the stroke required during large structural response (a maximum length of 30 cm is recommended (Mulligan, 2007)). Moreover, Mulligan (2007) recommended a piston diameter smaller than 20 cm, that is, D # 20 cm. An empirical relation to approximate the required damper stiffness was proposed by Rodgers et al. (2007), as follows: kd 5 6:29

ζd k

(5.237)

where k denotes the structural stiffness of SDOF system. Fig. 5.35 shows several curves for selection of piston diameter and chamber length, each one is associated with different device effective stiffnesses (Chase et al., 2006). The range of practical values of these parameters is also indicated as black rectangle in the figure. Hence, appropriate values can be selected with the use of estimated stiffness. More realistic experiments may be required for manufacturing these devices for application in real structures (see Step 9, Section 5.4.2.9). Design of semiactive-tuned mass dampers. This system is developed based on the conventional TMD system in such a way the tuned mass is an upper part of building itself (see Fig. 5.36A). To avoid excessive lateral stroke of the tuned mass (upper portion of building), the isolators (e.g., rubber or elastomeric bearings) are combined with a viscous damper (to get a passive TMD) or a resettable device (to get the SATMD), similar to what is shown in Fig. 5.36. Such a controlled building is idealized with an equivalent SDOF system (i.e., the lower part of building isolated by dampers from the upper portion) equipped with SA TMD (see Fig. 5.36B), generally resulting in a 2-DOF system. The overall effectiveness of such a system is based on the amount of seismic-induced energy transferred to the upper part (i.e., tuned mass), the size of the tuned mass,

391

CHAPTER 5 Design procedures for tall buildings

0.700

250 kN/m

0.600 0.500 Diameter (m)

392

125 kN/m

0.400 50 kN/m

0.300

25 kN/m 0.200 0.100 0.000 0.000

0.100

0.200 L0 (m)

0.300

0.400

FIGURE 5.35 Curves relating chamber length L0 to diameter D for different device nominal stiffness values. The black box indicates the range of possible design (Chase et al., 2006).

FIGURE 5.36 (A) Schematic of SATMD model using a resettable device and (B) idealized SDOF system with an SATMD. PTMD, Passive TMD. Adapted from Chey, M.-H., Chase, J.G., Mander, J.B., Carr, A.J., 2010. Semi-active tuned mass damper building systems: design. Earthquake Eng. Struct. Dynam. 39, 119
139.

5.4 Active, Semiactive, and Hybrid Systems

and the capability of resettable device in dissipating that energy. The sequence of parameters to define is as follows: •

•

•

•

Selection of a mass ratio μ The mass ratio is the ratio between the total mass of the isolated stories and the total mass of lower stories (stories located under the semiactive damper level). A large mass ratio can be selected (e.g., μ 5 0:5 (Chey et al., 2010)). Calculation of optimal frequency ratio and damping ratio of TMD The optimal tuning frequency, fopt , and supplemental damping ratio, ζ d;opt , can be calculated using expressions described in Appendix A (Tables A.1 and A.2), depending upon the type and position of the external excitation applied on the primary (SDOF) system. Calculation of optimal stiffness and damping Having determined the optimal frequency ratio and damping ratio, the optimal stiffness, kd , and damping, cd , coefficients associated with the TMD can be calculated as shown in Eqs. (5.113) and (5.114). Note that the calculated kd and cd are assigned, respectively, to rubber bearing and viscous damper, if a passive TMD is employed at interface of tuned mass and main mass. Since the scope of this procedure is to design the SA TMD, the kd should be considered as the sum of stiffness of semiactive (resettable) device (kdðresÞ ) and rubber bearings (kdðRBÞ ) (see Fig. 5.36B). Perform a time-history analysis Chey et al. (2010) suggest to model the 2-DOF system (Fig. 5.36B) and the spring member, kd 5 kdðRBÞ 1 kdðresÞ , using adequate software. For an effective controller evaluation, it is recommended (Chey et al., 2010) to perform history analysis using 30 earthquake time histories having multilevel hazard intensities (50% in 50 years (low intensity), 10% in 50 years (medium intensity), and 2% in 50 years (high intensity)); each of these three intensity levels may contain 10 pairs of earthquake records. Semiactive resettable devices can be simulated with Ruaumoko (Carr, 2007), where two types of relevant hysteretic loops are shown in Fig. 5.37 with (Fig. 5.37A) and without saturation (Fig. 5.37B). In the first type, the force in the device is proportional to the displacement until it reaches yield forces of the resettable device (i.e., saturation force); see Fy1 and Fy2 in Fig. 5.37. After it yields, the system behavior is perfectly plastic. For the type without saturation (Fig. 5.37B), the force is always proportional to the displacement with the device stiffness, and the system behavior can be defined with an idealized Bouc
Wen model (Wen, 1976) (see Chapter 4, Section 4.1.1.2.1, for more details about this model). In both types, the device is resetting once the force automatically drops to zero (at maximum and minimum displacements). It is recommended to set a maximum force of device equivalent to 13.8% of total structural weight multiplied by mass ratio (Hunt, 2002; Chey et al., 2010).

393

394

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.37 Hysteresis behavior of semiactive resettable device: (A) with saturation and (B) without saturation (Chey et al., 2010).

•

Note that in this design procedure of SATMDs, the main focus is on the stiffness of the device to be compared with passive TMD approaches. For the accurate modeling of dynamic characteristics (hysteretic loops) of such devices, experimental investigations are required as discussed in Chapter 7. Statistical assessment Having analyzed (in the previous step) the 2-DOF system with a suite of random seismic excitation, a statistical assessment of maximum structural responses (response spectra) can be carried out in this step. To this end, the use of lognormal statistics is recommended (Chey et al., 2010). Hence, the results obtained from the analysis of each earthquake suite (the set of 10 recorded histories) can be combined using two statistical parameters (median x^ and dispersion factor β) for each response (e.g., peak interstory drift, peak displacement, or peak acceleration) (Rodgers et al., 2007): ! n 1X x^ 5 exp lnðxi Þ n i51 sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi n   2 1 X β5 ln xi =x^ n 2 1 i51

(5.238)

(5.239)

where n is the number of samples (depending on the number of earthquake records) and xi is the maximum (spectral) response (displacement or acceleration) associated with ith sample (earthquake record).

5.4.2.9 Step 9: Verification analyses and quality control After designing the damping systems, it is important to check if the system satisfies the selected PO. In case this is not verified the device design parameters should be revised to improve the system performance. The explicit modeling of

5.4 Active, Semiactive, and Hybrid Systems

active-based damping systems is a difficult task and the detailed review of this is outside the scope of this publication. Thus, analytical expressions (models) and experimental tests are usually proposed for the response analysis of such systems (see Chapter 7, Section 7.2.2).

5.4.2.9.1 Active tuned mass dampers In this section, the controlled structural responses obtained on the basis of design approaches (Section 5.4.2.8.1) (Approaches 1
3) are presented. Consequently, the designer can compare the controlled and uncontrolled responses in order to assess the performance of active controlling system designed. Approach 1. Determination of the standard deviation of roof displacement. Since the PI of this design approach was based on the roof displacement, in this step, the relative standard deviation, σs , is given by (Ankireddi and Yang, 1996): ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s ffi πSb N σ s 5 φ1 ð H Þ m21 D

(5.240)

where N5

 2    B0 ðA2 A3 2 A1 A4 Þ 1 A3 B21 2 2B0 B2 1 A1 B22 A0   D 5 A1 ðA2 A3 2 A1 A4 Þ 2 A0 A23

(5.241) (5.242)

with B2 5 μ 1 μg ;B1 5 εgopt ;B0 5 ψgopt

8 > > < A4 5 μ 1 μ g 1 μμg ;A3 5 ð1 1 μÞεgopt 1 2ζ 1 ω1 μ 1 μg A2 5 μ 1 μg ω21 1 ð1 1 μÞψgopt 1 2ζ 1 ω1 εgopt > > : A1 5 ω21 εgopt 1 2ζ 1 ω1 ;A0 5 ω21

(5.243)

(5.244)

In Eq. (5.240), Sb is the uniform power spectral density of the stationary white-noise disturbance of wind type. This parameter can be simply calculated for the existing excitation using available routines in MATLAB software (Samui et al., 2016). Alternatively, for wind excitation, the nondimensional power spectral density can be given by (Balendra et al., 1995; Yan et al., 1998): 2

Sb 5

4k0 F Φ ωω1

(5.245)

where ω is the frequency of the harmonic excitation, and:

!αp 8 2 U r π4 Ψω1 U10 > > z > Φ5 h ; Uz 5 U10 10 i5=6 ; U r 5 > > Uz < 21 ðπ3 Ψω1 Þ2 > > ω1 lx F0 π4 > > Ψ5 4 ; F 5 2 ; F0 5 ρ0 A0 C0 Uz2 > : 2π U10 ω 1 m1

(5.246)

395

396

CHAPTER 5 Design procedures for tall buildings

where k0 is the ground surface drag coefficient; lx is the wave length; z is the height of structure; U10 is the mean wind speed at the reference height of 10 m; αp is the power law exponent; ρ0 is the air density; A0 is the structure frontal area; C0 is the drag coefficient; and Uz is the mean wind speed at the height z. Interested readers should refer to Choi and Kanda (1993) for the across-wind formulation of the nondimensional spectra of the wind force. It is worth mentioning that since the device stroke is dominated by the damper mass, the use of active control would increase the stroke. Hence, the device stroke should be checked to be in a reasonable range (Yan et al., 1999). Approach 2. Determination of mean-square responses. Given that the DPI case represented in Approach 2 is more practical than the VPI one, the normalized mean-square response of displacement and velocity of the main mass ; x_1 ) and  22  (x  122 22 _ _ , E x , E x auxiliary mass (x ; x ), respectively, indicated by E x d d 1 d , and 1   E x_ 22 , for both wind and earthquake excitations, can be computed as follows d (Nagashima, 2001): •

For wind excitation pffiffiffi pffiffiffiffiffiffi m ω2 ðμ 1 4ζ 2 Þ G1 1 d 1 3μ s  22  3 2ω1 md qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi E x1 5 pffiffiffi 4 μð1 1 μÞ G 1 m ω2 ð11μÞ21  G 1 2ζ 2 m ω2 μ21 1 d 1 1 d 1 s

E



x_ 22 1



pffiffiffi pffiffiffiffiffiffi 2  m ω2 ½4ð1 1 μÞζ 2 1 ðμ3 1 3μ2 1 μÞ G1 1 d 1 μð1 1 μÞðsμ2 1 2μ 1 3Þ 2ω1 md μ 1 2μ 1 3 5 pffiffiffi  qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4 μð11μÞ2 G1 1 md ω21 ð11μÞ21 G1 1 2ζ s 2 md ω21 μ21

E



x22 d



md ω21 md 2 ω41 pffiffiffi 2 2ð11μÞ3 G1 1 μð1 1 μÞ G1 1 μð11μÞ2 5 3 3=2 pffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi μ 4ω1 ma G 1 2ζ 2 m ω2 μ21 1

  5 E x_ 22 d

d

s

d

(5.248)

(5.249)

1

pffiffiffi 2ð1 1 μÞ G1 1 md ω21 ffi pffiffiffiffiffiffi 3=2 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4ω1 md μ G 1 2ζ 2 m ω2 μ21 1

•

s

(5.247)

(5.250)

1

For earthquake excitation pffiffiffi m ω2 ½4ð1 1 μÞζ 2 1 μ pffiffiffiffiffiffi G1 1 d 1 3μð1 1 μÞs  22  3 2ω1 ð1 1 μÞ md E x1 5 pffiffiffi  qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4 μ G1 1 md ω21 ð11μÞ21 G1 1 2ζ s 2 md ω21 μ21

(5.251)

pffiffiffi pffiffiffiffiffiffi m ω2 ½4ð1 1 μÞζ 2 1 μ G1 1 d 1 3μð1 1 μÞs  22  3 2ω1 md E x_ 1 5 pffiffiffi  qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4 μ G2 1 md ω21 ð11μÞ21 G1 1 2ζ s 2 md ω21 μ21

(5.252)

5.4 Active, Semiactive, and Hybrid Systems

  5 E x22 d

E



x_ 22 d



h ih i md ω21 md ω21 pffiffiffi G1 1 1 1 μ 2ð11μÞ5 G1 1 μ pffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4ω31 ma 3=2 μ G1 1 2ζ s 2 md ω21 μ21

(5.253)

pffiffiffi md ω21 G1 1 1 1 2ð11μÞ3 μ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 5 pffiffiffiffiffiffi 4ω1 ma μ3=2 G 1 2ζ 2 m ω2 μ21 1

s

d

(5.254)

1

The above expressions give the controlled responses in case the displacement feedback parameter G1 is already known. For evaluation purposes, if G1 5 0 is substituted into the above equations, the passive-controlled responses are comparable to those obtained from the active-controlled system. Approach 3. Determination of responses. Regarding Approach 3, the expressions of response analysis are represented for the system under both harmonic and random excitations, as follows: •

For harmonic excitation The displacements of main structure and ATMD are expressed, respectively, by: y1 ðtÞ 5 Aeiðωt2θ1 Þ xd ðtÞ 5 Be

(5.255)

iðωt2θ2 Þ

(5.256)

where θ1 , θ2, A, and B are given in Section 5.4.2.8.1. Consequently, the control force is given by:  uðtÞ 5 2md

•

G1 2 G3 G5 ω y1 1 G2 ω2d xd 1 2 ζ 1 ω1 y_ 1 1 2G4 ζ s ωd x_ d 1 y€ 1 G6 x€ d μ 1 μ μ 1

(5.257)

For random excitation Based on the proposed approach, the standard deviation of structural displacement, controlled by the ATMD, can be calculated by: rffiffiffiffiffiffiffiffiffiffi B4 π σs 5 a4 A4

(5.258)

where A4 and B4 are determined in Section 5.4.2.8.1. σs can be compared with the uncontrolled (inactive or passive) responses in order to check the system performance; for this purpose, the gain parameters G1
G6 must be set to zero in the definitions of A4 and B4 (Eqs. 5.213 and 5.214).

5.4.2.9.2 Semiactive (magnetorheological) dampers In order to evaluate the properties of a manufactured MR damper, conducting performance tests (magnetic field tests and dynamic tests) is appropriate (Xu et al., 2012) (see Chapter 7, Section 7.2.2.3 for more details). Moreover, the accuracy of

397

398

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.38 Bouc
Wen model of the MR damper.

the designed parameters of the device can be examined by conducting a finite element simulation on the magnetic circuit (Yang et al., 2014). Simpler dynamic models of MR dampers are proposed for the simulation of damper behavior and structural vibration control (Gamato and Filiskom, 1991; Spencer et al., 1997; Wereley et al., 1998; Yang et al., 2004). Based on the test data, Spencer et al. (1997) proposed to use a Bouc
Wen model (Wen, 1976) for MR dampers (Fig. 5.38, see Chapter 4, Section 4.1.1.2.1, for more details about this model). The force in the model is provided by the following relationship: F 5 c0 x_ 1 k0 ðx 2 x0 Þ 1 αz

(5.259)

where c0 is the damping coefficient; k0 is the linear spring coefficient; x and x_ are the absolute displacement, α is the postyield ratio and velocity; z is the evolutionary variable governed by: z_ 5 2 γ jx_ jzjzjn21 2 β x_ jzjn 1 Ax_

(5.260)

The parameters, γ, β, n, and A, are dimensionless quantities controlling the hysteretic behavior of the model and they should be identified by experimental tests for the designed MR device (Wen, 1976). Another model (see Fig. 5.39) was developed by Yang et al. (2004) in which the fluid inertial effect is represented by an equivalent mass, m; the accumulator spring stiffness, k; frictional force due to the damper seals, f0 ; and the postyield _ This velocity-dependent damping can be plastic damping coefficient, cðxÞ. expressed as follows: _ 5 a1 e2ða2 jx_ jÞ cðxÞ

p

(5.261)

which is defined as a mono-decreasing function with respect to the absolute _ The coefficients a1 , a2 , and p are positive and can be estimated using velocity x. experimental results (Yang et al., 2004).

5.4 Active, Semiactive, and Hybrid Systems

FIGURE 5.39 Proposed mechanical model of MR dampers.

Subsequently, the damper force can be estimated with the following relationship: f 5 mx€ 1 cðx_ Þx_ 1 kx 1 αz 1 f0

(5.262)

with the evolutionary variable z determined as shown in Eq. (5.260).

5.4.2.9.3 Semiactive viscous dampers For the verification of designed semiactive viscous damper, experimental tests should be conducted in order to establish the dynamic characteristics under various inputs. Moreover, the efficiency of the control law and hysteretic loop of device in adding supplemental damping can be investigated (Franco-Anaya et al., 2007). Furthermore, with respect to the size and force capacity of these devices, they are appropriate for conducting large-scale structural experimental testing such that results can be applicable for real structures. It is suggested to consulate the manufacturer of the prototype devices during the design process (e.g., hardware selection and piston and cylinder sizes) and relative tests (Mulligan, 2007). For more details, it is encouraged to refer to the experimental parts of research works by Mulligan (2007) and Mulligan et al. (2009) and also to Chapter 7, Section 7.2.2.3.

5.4.2.9.4 Semiactive-tuned mass dampers In this step, two sets of verifications are conducted on the SATMD. The first case is to verify the results of statistical assessment (Section 5.4.2.8.2), which is based

399

400

CHAPTER 5 Design procedures for tall buildings

on the analysis of 2-DOF system, using corresponding results computed by nonlinear analysis of a full model (MDOF system) of building structure and explicit modeling of SATMD (Chey et al., 2010). For this aim, Ruaumoko (2004) can be employed. In parallel to this, the experimental validation of semiactive device can be performed using devices testing (Mulligan, 2007; Mulligan et al., 2009) and reduced-scale test of structure equipped with semiactive device. The concept of experimental tests for such resettable devices is similar to those explained in the previous section (see Chapter 7, Section 7.2.2.3, for more details).

5.5 RETROFIT OF EXISTING BUILDINGS “Retrofit” means to improve structural capabilities (e.g., strength, stiffness, ductility, regularity in height/plan, structural integrity, and stability) of a building that shows deficiency, based usually on code requirements. Nevertheless, “seismic retrofit” or “seismic rehabilitation” can be defined as the design of measures to enhance the seismic performance of structural or nonstructural components of a building by refining deficiencies identified with a seismic evaluation in proportion to a selected PO (ASCE, 2017a). Similarly, wind retrofit means to enhance the building performance from a wind deficiency point of view. Retrofit code requirements are mainly focused on the seismic aspects and for this reason seismic provisions are mainly reviewed in the following section.

5.5.1 CODE REQUIREMENTS Few standards are available worldwide for building retrofit as shown in Section 5.1.1. In the following, the history of the development of the US standards is briefly reviewed since, based on the authors knowledge, it is considered the most comprehensive standard available. Other provisions (e.g., Japanese and Chinese) are briefly reviewed as well: •

United States: one of the first codes, which explicitly addressed the seismic evaluation methodology of existing buildings, was ATC-14 (ATC, 1987). This standard presented a method to guide practitioners in assessing existing buildings to identify potential seismic deficiencies and vulnerabilities. Moreover, recommendations for upgrading buildings to achieve the minimum life-safety PO were given. The procedures suggested in ATC-14 (ATC, 1987) are the extended version of analysis methods and evaluation concepts developed in ATC-3 (ATC, 1978) and ATC-6-2 (ATC, 1983). ATC-14 (ATC, 1987), a handbook for the seismic evaluation of existing buildings, was developed by FEMA (1992a). This standard, called FEMA-178 (FEMA, 1992a), sets a series of evaluation statements for zones with high seismicity. Moreover, FEMA-172 (FEMA, 1992b) developed further information on techniques to lessen seismic deficiencies for various construction types.

5.5 Retrofit of Existing Buildings

•

•

Subsequently, FEMA 310 (FEMA, 1998a,b), later published as ASCE 31-03 (ASCE, 2003), provided a procedure with tiers for seismic evaluation of existing buildings in view of multiple levels of seismicity and two POs (life safety and immediate occupancy). In parallel, Vision 2000 (SEAOC, 1995) was developed to be used for the evaluation and upgrade of existing buildings. This was the first document that fully summarized performance-based methodologies considering seismic hazard, performance levels, and building use. Moreover, FEMA 273 (FEMA, 1997) Guidelines for the Seismic Rehabilitation of Buildings was developed as a base for future development of performance-based building codes. Some features such as simplified and systematic rehabilitation methods and the procedures for incorporating new technologies into rehabilitated structures were discussed. The evolution of this code was realized with FEMA 356 (ASCE, 2000). This standard was then published as ASCE 41-06 (ASCE, 2007), demonstrating the procedures to rehabilitate existing building based on achieving different performance levels. Combining the standards ASCE 31-03 (ASCE, 2003) and ASCE 41-06 (ASCE, 2007), a most updated version, ASCE 41-13 (ASCE, 2013), was developed for the seismic evaluation and retrofit of existing buildings (that was recently updated with the 2017 version (ASCE, 2017b)). ASCE 41-13 (ASCE, 2013) introduces an approach including three-tier evaluation steps for existing building and the relative retrofit strategies. Japan: The Japanese Building Disaster Prevention Association (TJBDPA, 2005) provides a technical manual for the seismic evaluation and retrofit of existing RC buildings. Takewaki et al. (2013a,b) state that also the document released by the Japanese Government (MLIT, 2011) proposed provision for the retrofit of existing high-rise buildings in Japan based on a set of simulated long-period ground motions. China: The Chinese code (GB50223, 2008) proposed four protection categories to improve seismic protection requirements: moderate, standard, emphasized, and particular protection. Correspondingly, different design schemes are adopted for each of these categories.

In alternative to code prescriptive procedures, in literature there are available several nonprescriptive procedures, such as: •

•

•

Martinez-Rodrigo and Romero (2003) proposed a numerical scheme for evaluation and retrofit of multistory building using linear and nonlinear viscous dampers. Accordingly, a 6-story steel frame was investigated to find the optimum retrofit scheme. Weng et al. (2012) developed a simplified (step-by-step) design procedure for seismic retrofit of RC buildings using viscous dampers and proved its effectiveness as seismic retrofit design for earthquake-damaged frames structures. Li et al. (2015) developed a step-by-step procedure for design and retrofit of buildings equipped with metallic structural fuses using damage-reduction spectrum method. They assessed the performance procedure for a 5-story frame using the buckling restrained bracings (BRBs) as the structural fuse.

401

402

CHAPTER 5 Design procedures for tall buildings

Having briefly discussed the major procedures for the retrofit of existing buildings with dynamic modification devices, in the following the proposed design procedure, mainly based on US standards, is reviewed.

5.5.2 EVALUATION PROCEDURES BASED ON ASCE 41-13 (ASCE, 2013) The building evaluation process is the first step for defining what is the required (if any) seismic retrofit procedure of a building. In ASCE 41-13 (ASCE, 2013), three evaluation procedures are generally proposed: Tier 1, Tier 2, and Tier 3, which are briefly described in following: •

•

Tier 1 (screening procedure) Tier 1 refers to an evaluation report checklist that allows to identify potential defects in a building based upon performance of similar buildings in past earthquakes. This method gives the opportunity to exert a quick evaluation of the building elements (structural, nonstructural, and foundation), geologic hazard, and site conditions. Two structural performance levels are allowed: immediate occupancy and life safety. While for the nonstructural elements the following are permitted: position retention and life safety. After performing Tier 1 checklist, if any defect is identified, then the designer may proceed to a more detailed evaluation method (Tier 2). Tier 2 (deficiency-based procedure) This method is applicable for certain building types and POs. It allows to establish if evaluated potential deficiencies (e.g., identified using Tier 1 method) need mitigation. In most of the cases, it might not be required to analyze the response of the entire structure. Therefore, for obtaining the building properties, it is permitted to use destructive and nondestructive examinations. The scope of this approach is to identify the cause and the extent of the deficiencies. Tier 2 evaluation procedure involves conducting simplified linear analyses (i.e., linear static and linear dynamic procedures) of the building in view of the potential defects identified in Tier 1. This tier is suitable for small, rather simple, buildings (e.g., those with not complex structural systems) and for buildings without a need to advance analytical approaches (e.g., those based upon nonlinear analysis techniques). This is because the common defects are almost well understood and the mitigation approaches are straightforwardly attainable. The structural performance levels (immediate occupancy and life safety) and nonstructural performance levels (position maintenance and life safety) are accepted for Tier 2 procedure. In case partial retrofit is necessary (e.g., due to limited funding or being done while the building is occupied), using Tier 2 procedure retrofitting should be adopted in a priority order, that is, the more critical ones at first. In alternative, the designer can proceed with Tier 3 for a more comprehensive evaluation. Indeed, Tier 2 procedures may produce more conservative results compared to those of Tier 3 because of a diversity of simplifying assumptions (ASCE, 2013).

5.5 Retrofit of Existing Buildings

•

Tier 3 (systematic procedure) Tier 3 is an evaluation procedure in which the response analysis of the building is fully performed under seismic hazards. This systematic evaluation procedure may be applied at any time (e.g., at the beginning of evaluation process) or may be used to further evaluate the potential deficiencies identified in Tier 1 or Tier 2 procedures. This kind of evaluation is more suitable for complex structures (e.g., tall buildings) and it requires to focus on the nonlinear structural response. Nonlinear analysis procedures may be employed in preliminary evaluations without testing given that the required testing is conducted before the retrofit. Albeit Tier 3 procedure is complicated and expensive to carry out, it frequently leads to construction savings. Starting the evaluation immediately with Tier 3 procedure is often suitable when the building has significant seismic defects regarding the selected PO (ASCE, 2013).

In the case of tall buildings, Lai et al. (2015) state that Tier 1 is not applicable, and Tier 3 is automatically required. According to ASCE 41-13 (ASCE, 2013), seismic isolations or damping systems are considered as possible retrofit technique and their analysis and design are included in Tier 3 procedure. Readers should refer to ASCE 41-13 (ASCE, 2013) (and the most updated version of ASCE 41-17 (ASCE, 2017b)) for more details about the limitations of using Tier 1 and Tier 2 based on the building type (load-resisting structural system) and number of stories.

5.5.3 STEP-BY-STEP PROCEDURE Based on the literature review presented in the previous sections, a step-by-step procedure for the seismic retrofit process of tall buildings is presented (see Fig. 5.40). This is mainly based on the requirements of ASCE 41-13 (ASCE, 2013) and the main focus is on supplementary damping (and isolation) systems as retrofit strategy, while other retrofit strategies useful to tall buildings are just addressed.

5.5.3.1 Step 1. Initial considerations The first important step involves a building evaluation to determine if the existing structure has acceptable performance capacity and to identify eventual deficiencies. According to ASCE 41-13 (ASCE, 2013), initial considerations may include: •

•

Structural characteristics of the building. The structural characteristics including the type of structural systems, type of material, and geometrical dimensions should be analyzed before starting retrofit design. Seismic hazards at the site. Seismic hazards other than ground shaking may exist at the building site. The risk and possible damage from geologic site

403

404

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.40 Step-by-step procedure for retrofit process of existing buildings.

• •

hazards should be considered. In some cases, it may be more feasible to mitigate the site hazard than retrofit the building. Results of prior seismic evaluations. If there is any seismic evaluation, performed previously, it could be considered before the current retrofit process. Building use and occupancy requirements. To estimate the significance of potential temporary or permanent disruptions associated with various

5.5 Retrofit of Existing Buildings

•

• •

•

risk-mitigation schemes, the use of the building must be considered. Regarding vulnerable buildings, the occupancy can be reduced; moreover, redundant facilities can be provided, and nonhistoric buildings can be demolished and replaced. Historic status of the building. The historic status of the building should be determined if it is at least 50 years old. This should be made as early as possible, since it could influence the choices of retrofit strategy. Economic considerations. The range of costs and impacts of retrofit (e.g., the variation associated with different POs) should be considered. Societal issues. Since existing buildings have been built with the use of earlier standards and often are occupied, several issues may arise, such as distribution impacts on various parts of the community, reduced business interruption, occupancy dislocation and the loss of housing, the treatment of historic properties, and methods for financing seismic rehabilitation (FEMA 275 (FEMA, 1998a,b)). Local jurisdictional requirements. Particular local jurisdiction requirements shall be carefully evaluated since they could be more stringent than the main standard.

5.5.3.2 Step 2. Performance objectives and hazard levels After the first step of initial evaluation, it is important to define the target PO as described by ASCE 41-13 (ASCE, 2013). A PO may consist of one or more combinations of selected “seismic hazard levels,” with a target “structural performance level” and a target “nonstructural performance level.” Performances can be defined qualitatively based on different aspects, such as safety of building occupants during and after the event (earthquake); cost and possibility of reverting the building to its preevent condition; period required for the building to be converted from service to effect repairs; and influences of economical, architectural, or historical problems on the larger community. These features are directly related to the level of damage that the building and its consisting systems would withstand during an event (ASCE, 2013). In other words, the PO, selected as a basis for design, specifies the cost and practicality of a project, and the achieved advantages in terms of enhanced safety, decrease in damage, and interruption of use in case of future earthquakes. To select the appropriate level, Table 5.28 lists the POs for a typical building. Based on the code recommendations three different levels of POs can be selected: • • •

Basic POs for existing building (BPOE) Enhanced PO Limited PO To reach each PO the following objectives need to be reached:

•

BPOE:
g and i

405

406

CHAPTER 5 Design procedures for tall buildings

Table 5.28 Performance Objectives (ASCE, 2013) Target Building Performance Levels

Seismic Hazard Level 50%/50 years BSE-1E (20%/ 50 years) BSE-2E (5%/ 50 years) BSE-2N (ASCE 7-16 (ASCE, 2017a) MCER)

•

•

Operational Performance Level (1-A)

Immediate Occupancy Performance Level (1-B)

Life Safety Performance Level (3-C)

Collapse Prevention Performance Level (5-D)

a e

b f

c g

d h

i

j

k

l

m

n

o

p

Enhanced POs:
g and i, j, m, n, o, or p
l and e or f
g and l and a, or b
k, m, n, or o alone Limited POs:
g alone
l alone
c, d, e, or f

To use this table, two factors should be defined: seismic hazard level (Section 5.5.3.2.1) and target building performance level (Section 5.5.3.2.2). Note that another PO, called basic PO equivalent to new building standards (BPON), is defined in ASCE 41-13 (ASCE, 2013). The scope of this objective is to retrofit the building such that it reaches the same performance level of a new structure based on the new building standard requirements.

5.5.3.2.1 Seismic hazard level Seismic hazard due to ground shaking can be defined as acceleration response spectra or ground motion acceleration histories specified based upon a probabilistic or a deterministic analysis (ASCE, 2013). This hazard may depend upon the location of the building with respect to faults, the regional and site-specific geologic and geotechnical features, and the specified seismic hazard levels. Different seismic hazard level can be defined (ASCE, 2013) and relative 5% damped acceleration response spectrum for short period (Ts 5 0:2 second), SS , and long period

5.5 Retrofit of Existing Buildings

(T1 5 1 second), S1 , in the maximum direction of horizontal response, can be determined as follows: •

•

•

•

Basic safety earthquake-2 (BSE-2N). It is equivalent to MCER to be used for the BPON standards. Therefore, it is computed using values of SS and S1 taken from the MCER spectral response acceleration contour maps (ASCE, 2017a). Basic safety earthquake-1 (BSE-1N). Defined as two-thirds of the BSE-2N useful for the BPON standards. Therefore, it is estimated using two-thirds of the SS and S1 values obtained for the BSE-2N seismic hazard level. Basic safety earthquake-1 (BSE-1E). It is equivalent to a seismic hazard with a 20% probability of exceedance in 50 years (lower than the BSE-1N). It is computed using values from approved 20%/50-year maximum-direction spectral response acceleration contour maps (SS and S1 ). Values for BSE-1E are not required to be greater than those achieved for BSE-1N. Basic safety earthquake-2 (BSE-2E). It is considered as a seismic hazard with a 5% probability of exceedance in 50 years (lower than the BSE-2N). It is computed using values from approved 5%/50-year maximum-direction spectral response acceleration contour maps (SS and S1 ). Values for BSE-2E are not required to be greater than those obtained for BSE-2N.

In addition to the above-mentioned procedures, it is also possible to specify the acceleration response spectra with the use of site-specific procedure. This is based on the geologic, seismologic, and soil characteristics associated with the building site. Reader may refer to Section 2.4.2 of ASCE 41-13 (ASCE, 2013) and also Section 5.2.1.1 for more details. Independent ground motion acceleration histories with magnitude, fault distances, and source mechanisms can be chosen for the BSE-1N, BSE-2N, BSE-1E, or BSE-2E seismic hazard levels. At least three data sets of ground motion acceleration, for site seismic hazard, should be used for the RHA, where each one includes two horizontal components. In case the vertical motion is important, two horizontal components and one vertical component of at least three records should be selected and scaled. It is recommended to refer to Section 2.4.2.2 of ASCE 4113 (ASCE, 2013) or to Section 16.2 of FEMA P-1050 (FEMA, 2015) for a more detailed discussion (as well as Section 5.2.1.1). The required number of ground motion pairs (N) is listed in Table 5.29 depending on the building site distance from the active fault. These limiting numbers are for POs of existing buildings (e.g., BPOE).

5.5.3.2.2 Target building performance level A target building performance level may be composed by structural and nonstructural target performance levels. Fig. 5.41 illustrates estimated target performance levels and ranges in buildings (ASCE, 2013). With regard to structural and

407

408

CHAPTER 5 Design procedures for tall buildings

Table 5.29 Required Number of Ground Motion Acceleration Records and Method of Response Analysis (ASCE, 2013) Building Site Distance From Active Fault (km)

Method of Calculating Responses

Limiting Number of Earthquake Record Pairs (N)

.5 .5 #5 #5

Average Maximum Average Maximum

N $ 10 3#N#9 N $ 10 3#N#9

FIGURE 5.41 Target building performance levels and ranges. Adapted from ASCE, 2013. ASCE41-13: Seismic Evaluation and Retrofit of Existing Buildings. American Society of Civil Engineers, Reston, VA.

nonstructural performance levels, the damage patterns of structural systems and elements can be found in Section 2.3 of ASCE 41-13 (ASCE, 2013). The basic POs for existing buildings depend upon the risk category defined in ASCE 41-13 (ASCE, 2013), as given in Table 5.30, given the risk category, hazard level, and the evaluation procedures (Tiers 1
3).

5.5 Retrofit of Existing Buildings

Table 5.30 Basic POs for Existing Buildings (ASCE, 2013) Evaluation Procedure

Tier 3

Hazard Level Risk Category I and II

III

IV

BSE-1E LS structural performance LS nonstructural performance (3-C) DC structural performance PR nonstructural performance (2-B) IO structural performance PR nonstructural performance (1-B)

BSE-2E CP structural performance Nonstructural performance not considered (5-D) LS structural performance Nonstructural performance not considered (4-D) LS structural performance LS nonstructural performance not considered (3-D)

Note: LS, life safety; LS, limited safety; IO, immediate occupancy; CP, collapse prevention; DC, damage control; PR, position retention.

Enhanced PO can be achieved if one of the following options or a combination of them is considered: • • •

Target building performance levels are higher than those of BPOE. Seismic hazard levels are higher than those of BPOE. Risk categories are higher than those of BPOE.

Accordingly, some possible enhanced POs can be the following, as recommended by ASCE (2013): •

• •

For a certain risk category, target structural or nonstructural performance levels which are higher than those of the BPOE at the hazard level of BSE1E, or BSE-2E, or both of them. For a certain risk category, target structural or nonstructural performance levels of the BPOE at a hazard level higher than BSE-1E, or BSE-2E, or both of them. Target building performance levels of the BPOE using a risk category higher than the one of the building.

If a PO is less than that of the BPOE, then a limited PO is obtained. This may be achieved by using a reduced PO or a partial retrofit objective. The reduced objective is defined when a lower seismic hazard level or a lower target building performance level than that of the BPOE is applied. The partial objective refers to the retrofit of parts of the building without the complete rehabilitation of the resisting system. It is worth mentioning that the use of a limited PO should satisfy the following requirements (ASCE, 2013): •

It should not lead to any reduction in the structural or nonstructural performance levels of the existing building for the same seismic hazard level.

409

410

CHAPTER 5 Design procedures for tall buildings

• •

It should not result in a new structural irregularity or in more severe structural irregularity. It should not lead to an amplification of the seismic forces in any component which is potentially incapable of resisting such forces.

5.5.3.3 Step 3. Evaluation of existing building Before starting the retrofit process of a building, it is recommended to obtain its as-built information to evaluate any deficiency. All the components of the building should be evaluated and analyzed so that the selected PO is satisfied. An appropriate evaluation procedure can be selected depending on the selected PO, level of seismicity, and building type. Based on ASCE 41-13 (ASCE, 2013) requirements, Tier 3 procedure is selected for the evaluation of tall buildings. The required investigations/information to be collected on the building are as follows: •

•

• •

•

Building type. Determined based on the seismic load-resisting system and the diaphragm type (see Table 3-1 of ASCE 41-13 (ASCE, 2013)). If the structural system of the building along each direction is different, separate building types may be chosen. Building configuration. Related to the type and arrangement of existing structural components of the vertical and seismic force-resisting systems as well as of the nonstructural elements with effective stiffness/strength when bearing loads. Potential seismic deficiencies, for example, discontinuities in the load path, weak links, irregularities, and insufficient strength/deformation capability in components, may be identified with the use of building configuration information. Component properties. Needed to analyze the strength and deformation capacities of structural components (e.g., beams, columns, and diaphragms). Site and foundation information. Determined based on existing building documentation, visual site inspection, or an investigation program of sitespecific subsurface (in case of inadequate available data to compute foundation capacities or to realize the presence of geologic site hazards). Adjacent buildings’ information. These are important to understand some of the possible phenomenon:
Building pounding. It can change the basic response of the building to earthquake, imposing extra inertial loads and energy to the building from adjacent ones. For example, extreme local damage to structural elements at the zones of impact may occur, especially where the floor and roof levels of adjacent buildings are not aligned in height.
Shared element condition. Shared elements (e.g., party walls) between existing building and adjacent ones may create two problems depending on if the buildings move independently or as an integrated unit. In the first case, a partial collapse may occur at the position of shared element.

5.5 Retrofit of Existing Buildings




Concerning the second case, an extreme load may be imposed to the buildings due to the presence of additional mass and inertial forces. Hazards from adjacent building due to potential falling debris, rooftop equipment, fire, and blast pressures. Refer to Section 3.2.5.3 of ASCE 4113 (ASCE, 2013) for further details.

Three levels of data collection knowledge are presented in ASCE 41-13 (ASCE, 2013): minimum data collection, usual data collection, and comprehensive data collection (Table 5.31). In this table, a knowledge factor, κ, is used to account for uncertainty in the collection of as-built data. This factor is used to express the confidence with which the properties of the building components are known, where calculating component capacities. It is worth mentioning that if the material testing is delayed up to retrofit construction, this may potentially lead to reevaluation or redesign of the retrofit due to differences between the assumed material properties and those determined by testing (ASCE, 2013). After performing the evaluation process, an evaluation report could be prepared accordingly. This report may contain, at the least, the following details (ASCE, 2013): • •

•

Scope and intent. The aim of the evaluation including a summary of the evaluation procedure(s) adopted and level of investigation applied. Site and building data. This might include: • General building description, including number of stories and dimensions, availability of original design, and construction documents • Structural system description, including material properties, load-resisting systems, floor diaphragms, basement, and foundation system • Nonstructural system description, including those affecting seismic performance of the building or being dangerous • Common building type and occupancy • Performance level • Level of seismicity • Soil type and conditions Findings. A list of seismic deficiencies identified could be prepared.

For example, a list of common seismic deficiencies identified in existing tall buildings previously built in Los Angeles, Oakland, and San Francisco are as follows (Lai et al., 2015): • • • • • • •

Low design base shear was considered at the design time. Small magnitude of drift limits were considered during the design. Simplified (static) analysis methods, which does not account for dynamic features, were used for the initial design. Vertical irregularities (especially close to the base of tall buildings) and plan irregularities. The case “strong-column-weak-beam” was not checked for member sizing. Beam and column joints and panel zones were not seismically designed. Column forces were not appropriately transferred to foundations.

411

Table 5.31 Data Collection Requirements (ASCE, 2013) Data Performance level Analysis procedures Testing

Level of Knowledge Minimum

Usual

Comprehensive

Life safety or lower

Life safety or lower

Greater than life safety

Linear static, linear dynamic

All types

All types

No tests

Usual testing

Comprehensive testing

Drawings

Design drawings or equivalent

No drawings

Design drawings or equivalent

Construction documents or equivalent

Condition assessment

Visual

Visual

Comprehensive

Visual

Comprehensive

Visual

Comprehensive

Material properties

From default values

From design drawings

From default values

From drawings and tests

From usual tests

From documents and tests

From comprehensive tests

Knowledge factor (κ)

0.75

0.9

0.75

1.00

1.00

1.00

1.00

5.5 Retrofit of Existing Buildings

5.5.3.4 Step 4: Model and analyze existing building In this step, the designer should model the building load-resisting structural system using appropriate software. At this level, all existing structural members effective on the response should be explicitly modeled (see Section 5.1.3.2.1). The following points are useful to be considered in modeling (ASCE, 2013): • •

•

• • • •

The load
displacement features of all components of building should be directly included in modeling. In case that the Ritz vector
based nonlinear response analysis is carried out using commercial software, the following three requirements shall be satisfied (see Section 5.1.3.2.3): • Sufficient modes should be included in the analysis so that at least 90% of mass participation is captured. • Time steps should be chosen as small as possible to confidently attain solution convergence. • Adequate vectors should be considered to accurately get local dynamic response of elements. For the analysis based on the 2D model, the deformations and forces in building elements should be evaluated under selected scaled ground motions. While for the 3D case, they should be assessed under a suite (series) of ground motions randomly oriented if the building is in a site at least 5 km far from an active fault. If the site distance from the fault is less than 5 km, the fault-normal components should be used for the analysis (ASCE, 2013). Inherent damping can be accounted using Rayleigh damping (ASCE, 2013; Lai et al., 2015) (see Section 3.3.1 (Chapter 3) for further details). P-Δ effects should be included in the mathematical model of building. SSI effects can be modeled explicitly if the spectral accelerations increase due to increase in fundamental period (ASCE, 2013). Uplift effects due to earthquake on the tension part of an element should be included in the analytical model as nonlinear DOFs (ASCE, 2013).

It is noted that nonlinear static (pushover) analysis, according to ASCE 41-13 (ASCE, 2013), is not allowed to be used for the analysis of tall buildings (since it neglects higher mode effects) (Lai et al., 2015). However, its uses may give basic information about postyield performance of the building. Having built the mathematical model of bare building, time-history analyses should be conducted for each pair of (horizontal) ground motion records (as determined in Step 2 (Section 5.5.3.2)). This is recommended by ASCE 41-13 (ASCE, 2013) to minimize or eliminate unnecessary seismic retrofits, since deformation and forces in building components can be determined more accurately. The following recommendations are given by ASCE 41-13 (ASCE, 2013) while conducting nonlinear analysis: •

The results obtained from nonlinear (dynamic) analysis can be straightforwardly compared with those achieved using tests to evaluate the performance of building components under a selected ground motion.

413

414

CHAPTER 5 Design procedures for tall buildings

•

•

For the structural component responses that are independent of the direction of loading (e.g., shear about the same axis in a beam/column/wall, plastic hinge rotation in walls, and building drifts), the maximum response (force or displacement) is computed as the maximum absolute responses obtained under each history analysis, and the average response should be calculated as the mean of the maximum values. For structural component responses that are dependent on the direction of loading (e.g., axial tension versus compression in a column and bending moment in an asymmetrical RC beam), the maximum response is specified independently for each loading direction as the maximum positive and minimum negative responses under each history analysis; the average response should be calculated independently for each loading direction as the mean of maximum values.

Based on the analyses results obtained in this step, there may be potential deficiencies with respect to some structural responses and the relative acceptance criteria can be checked in the next step.

5.5.3.5 Step 5. Acceptance criteria ASCE 41-13 (ASCE, 2013) divides the different components as primary and secondary to define their acceptability actions (i.e., forces and deformation). Primary are components of the lateral force
resisting system and secondary are not components of the lateral force
resisting system but they are affected by the structure lateral deformation. In any case, the components shall be able to carry gravity loads. Each action shall be classified (Fig. 5.42 and Table 5.32) as follows: •

Deformation controlled (ductile): Type 1 when d $ 2g, Type 2 when e $ 2g for primary components or f $ 2g for secondary component, Type 3 for secondary component if f $ 2g (Fig. 5.42). Expected material strength, QCE (mean value of the component resistance), shall be used to define the curves.

FIGURE 5.42 Component force
deformation curve. Adapted from ASCE, 2013. ASCE41-13: Seismic Evaluation and Retrofit of Existing Buildings. American Society of Civil Engineers, Reston, VA.

5.5 Retrofit of Existing Buildings

Table 5.32 Example of Components Force- and Deformation-Controlled Actions for Framing Systems (ASCE, 2013) Component

Deformation-Controlled Action

Force-Controlled Action

Moment (M)

M, V

Shear (V) Axial load (P), V Va P

P

V P, V, Mb M, Vc


P P P, M P, V, M P, V, M

Moment Frame Beam Columns Joint Shear walls Brace Frame Braces Beams Columns Shear link Connections Diaphragms a

V could be a deformation-controlled action in steel moment frame. P, V, and M actions could be deformation controlled for steel and wood constructions. c M and V are force controlled for diaphrams that carries lateral loads from vertical force-resisting elements above the diaphram level. b

•

Force controlled (nonductile): in all the other cases in which deformationcontrolled criteria are not verified. Lower bound material strength, QCL (mean value minus one standard deviation of the component resistance), shall be used to define the curves.

The strength Q is taken as the expected strength (QCE ) for deformationcontrolled action and the lower bound estimate (QCL , defined as the mean minus one standard deviation) for force-controlled actions, (ASCE, 2013). These values are defined for different structural materials in ASCE (2013). For deformation-controlled action ASCE 41-13 (ASCE, 2013) defines a generalized force
deformation curve (Fig. 5.43) for component modeling and acceptance criteria. The behavior can be divided as follows: • • • •

Linear from point A to B (that is the yield point) Strain hardening from point B to C Strength degradation from point C to D Reduced strength from point D to E

This curve can be given in terms of deformation (displacement, rotation) or deformation ratio as shown in Fig. 5.43A and B, respectively. Values for defining the modeling curve are given in the standard and their review is outside the scope of this publication. Similarly, ASCE 41-13 (ASCE, 2013) provides acceptance criteria limits for different building performance levels (Fig. 5.43C). In addition, ASCE 41-13 (ASCE, 2013) defines deformation nonlinear capacities for different

415

416

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.43 Force
deformation modeling and acceptance criteria for deformation-controlled action as a function of (A) deformation, (B) deformation ratio, (C) and acceptance criteria. Adapted from ASCE, 2013. ASCE41-13: Seismic Evaluation and Retrofit of Existing Buildings. American Society of Civil Engineers, Reston, VA.

materials and structural components. Given the fact that nonlinear procedures shall be utilized (Section 5.5.3.4), the component capacity shall be calculated as given in Table 5.33.

5.5.3.5.1 Load combinations Different from new construction code requirements (Section 5.2.1.1.4 (ASCE, 2017a)), the load combinations for existing buildings are the following (ASCE, 2013): •

Linear procedures with deformation-controlled actions Seismic and gravity additive: 1:1ðD 1 L 1 SÞ 1 QE

(5.263)

Seismic and gravity counteracting: 0:9D 1 QE

•

(5.264)

Linear procedure with force-controlled actions Seismic and gravity additive: 1:1ðD 1 L 1 SÞ 1

QE C1 C2 J

(5.265)

5.5 Retrofit of Existing Buildings

Table 5.33 Nonlinear Procedure Component Capacity (ASCE, 2013) Parameter Deformation capacity (existing component) Deformation capacity (new component) Strength capacity (existing component) Strength capacity (new component)

Deformation-Controlled Action

Force-Controlled Action

κx Deformation limit




Deformation limit







κxQCL




QCL

κ is the knowledge factor as determined in Table 5.31.

Seismic and gravity counteracting: 0:9D 1

QE C1 C2 J

(5.266)

where J is the force reduction factor taken as the smallest demand capacity ratio for the component in the selected load path (greater than or equal to 1). In alternative, it can be assumed to be equal to 2.0, 1.5, and 1.0 for high, moderate, low level of seismicity, respectively. C1 is the modification factor to account for the maximum inelastic displacement in linear response and C2 is the factor to be taken into account the effect of the hysteresis loop (see ASCE (2013) for further consideration about these factors).

5.5.3.6 Step 6. Retrofit strategies After having identified the deficiencies in the existing building, selecting an appropriate retrofit strategy is required to achieve the target PO. The retrofit solution adopted should be considered in the building model (developed in Step 4, Section 5.5.3.4) through its effects on stiffness, strength, yield behavior, and deformability. Several retrofit strategies are prescribed in ASCE 41-13 (ASCE, 2013), such as: • • • • • •

Local modification of components Removal or reduction in existing irregularities Global structural stiffening Mass reduction Seismic isolation Supplementary energy dissipation

For large buildings (e.g., tall buildings), it is recommended to represent several retrofit strategies and compare alternative ways of removing deficiencies (ASCE, 2013). Note that using additional damping for retrofit purposes may be accompanied with the use of other retrofit strategies in order to further improve

417

418

CHAPTER 5 Design procedures for tall buildings

the building performance (Lai et al., 2015). Herein, retrofitting by means of dynamic modification devices is described.

5.5.3.6.1 Step 6.1: Selecting suitable type of dynamic modification devices In general, two types of dynamic modification devices (seismic isolation and distributed damping system, and no indication about mass damping system is given) have been used for retrofit aims (ASCE, 2013), which are as follows: •

•

Effectiveness of seismic isolation is higher for relatively stiff buildings with low profiles and large mass, and it is less effective for light and flexible structures. Guidelines about how to select an appropriate type of isolation system (bearing) are discussed in Section 5.3.1.5. Distributed damping system may be effective for the modification of defects of a building due to excessive deformations caused by insufficient global structural stiffness. Significant deformations (or strokes) are often required in dissipation devices to substantially dissipate energy. This is achievable where the structure experiences significant lateral displacements (i.e., relatively flexible structures with some inelastic deformation capacity). In general, the structural displacements are often reduced by the addition of damping systems, but the forces transferred to the structure may essentially be increased. See the step-by-step procedure for distributed dampers for a detailed discussion (Section 5.2.1). According to Lai et al. (2015), the common supplementary damping systems for retrofitting tall buildings in the United States are BRB, fluid viscous dampers, and viscous wall dampers. Based on some observations on postearthquake performances after major earthquakes (e.g., 2011 Tohoku of Japan and 2003 Colima of Mexico), it is concluded that supplementary viscous dampers in high-rise buildings satisfy the life-safety requirements; moreover, such devices are efficient in controlling damage level (EERI, 2012; Takewaki et al., 2013a,b).

5.5.3.6.2 Step 6.2: Determining appropriate configuration and distribution for dynamic modification system After selecting the type of dynamic modification system, the suitable configuration and distribution (in plan and in height) can be specified correspondingly. If the selected dissipating device is of distributed type, user is encouraged to refer to Section 5.2.1.6. For mass dissipating devices, user can refer to Section 5.2.2.7. Instead, for base isolation systems, one can refer to Section 5.3.1.5 for the choice of location and distribution of isolators.

5.5.3.6.3 Step 6.3: Select total target damping The inherent damping of existing building should be selected for all the modes, as recommended by ASCE 7-16 (ASCE, 2017a). The designer should refer to

5.5 Retrofit of Existing Buildings

Section 3.2 for further discussion on the choice of inherent damping. Moreover, a total damping required for the design purposes can be selected with respect to the dominant (wind-based or seismic-based) design type. Sections 5.2.1.5 (for distributed dampers), 5.2.2.6 (for mass dampers), and 5.3.1.4 (for base isolation systems) represent more details about the damping selection.

5.5.3.7 Step 7. Damping system preliminary design In this step, based upon the selected damping system and its distribution, a preliminary device design should be carried out. For distributed dampers (viscous, viscoelastic, friction, and hysteretic), Section 5.2.1.7 properly represents the preliminary determination of damping properties. When dealing with mass damping systems (TMD, TLD, and TLCD), Section 5.2.2.8 addresses the optimal characteristics and then mechanical properties of devices. Concerning the base isolation systems, Section 5.3.1.7 should be used for preliminary design. In the case of distributed dampers, the code requires that the device capacity is increased for the maximum values for design and rare earthquake occurrence, in order to sustain larger displacement and velocities, with the following details: •

•

When more than four energy dissipation devices per primary direction are utilized (with at least two located on each side of the center of stiffness), the devices should resist a demand greater than or equal to 130% and 200% of the maximum values for the rare earthquake and DE respectively. When less than four energy dissipation devices per primary direction are utilized (or fewer than two devices are located on each side of the center of stiffness), the devices should resist a demand greater than or equal to 200% of the maximum values for the rare earthquake.

The dissipation device’s nominal design properties shall be bounded based on lambda factors (λ), as shown in Section 5.1.3.3.3, which are related to the material property’s dependence on manufacturing tolerances, environmental effect, aging, etc.

5.5.3.8 Step 8. Update building model and perform analyses Having defined the retrofit strategy, the proposed design should be verified by analyzing the building with the adopted retrofit measures. To this end, first the mathematical model of bare building (constructed in Step 4, Section 5.5.3.4) can be updated by adding the explicit model of damping (or isolation) systems designed. For distributed dampers, the user can refer to Section 5.2.1.8, which shows explicit modeling of viscous, viscoelastic, and hysteretic dampers. In the case of isolated damping systems, Section 5.2.2.9 addresses the simulation approaches. Regarding base isolation systems, Section 5.3.1.8 introduces the modeling approach. For the analyses, both linear and nonlinear procedures can be utilized depending on different criteria. Linear analysis procedures can be utilized only when devices are present in all stories of the upgraded building, and both upper and

419

420

CHAPTER 5 Design procedures for tall buildings

lower bound property analyses should be performed. The criteria for the utilization of linear procedures are the following: essentially elastic behavior of the main structure; supplemental damping shall not exceed 30%; secant stiffness at maximum displacement of the dissipation devices shall be included in the model of the rehabilitate building; and dissipation devices shall be considered while evaluating building regularity. Both static and dynamic procedures are allowed. The static linear procedure can be utilized when the following requirements are satisfied: the ratio between the story resistance and demand shall range between 80% and 120% of the average ratio for all stories (to be satisfied by displacement-dependent devices only), with the resistance provided by the dissipation devices in a story required not to exceed 50% of the total story resistance. ASCE 41-13 (ASCE, 2013) requires the calculation of the design actions (for both upper and lower bound properties) at three distinct stages of deformation: (1) at maximum drift, (2) at maximum velocity and zero drift, and (3) at maximum floor acceleration. Dynamic linear procedure, such as response-spectrum procedure, can be utilized for both displacement- and velocity-dependent devices. Damping values shall be calculated independently for each mode of vibration and the demand calculated by dynamic analysis shall be more than 80% of the one calculated from static linear procedures. In case the linear procedures are not allowed, nonlinear procedures shall be utilized. Nonlinear procedures shall include the nonlinear force
displacement characteristics of the dynamic modification devices, as well as their dependency on excitation frequency, temperature, deformation, velocity, etc. Therefore, this requires multiple analyses with bounded properties’ values to consider these effects (Section 5.1.2.4). ASCE 41-13 (ASCE, 2013) also states the importance of taking accidental eccentric in dynamic analyses through multiple analyses with mass shifting or with amplification factors on forces, drifts, and deformations into consideration. From the analyses results, the desired structural responses of retrofitted building (e.g., story shear, story drift, additional equivalent damping ratio) can be obtained for checking acceptance criteria (see Section 5.5.3.5) or for comparison with those obtained from bare building analysis.

5.5.3.9 Step 9. Check retrofit response acceptability In this step, response of the bare building (Step 4, Section 5.5.3.4) can be compared with that obtained from the analysis of retrofitted building (Step 8, Section 5.5.3.8). This provides the opportunity to see the variation in the structural responses under different hazard levels and load combinations. For instance, Weng et al. (2012) recommended the comparison of story shear forces and story drifts between bare building and retrofitted building when viscous dampers are utilized for retrofitting framed buildings. In general, comparing various responses of bare-base and upgraded-base responses helps designer realize which response is mitigated and how much is the level of mitigation reached.

5.5 Retrofit of Existing Buildings

The selected retrofit measures (e.g., supplementary damping system) are satisfied if the acceptance criteria (Section 5.5.3.5) for the selected PO selected (Step 2, Section 5.5.3.2) are met for all the building’s element. In case this is not satisfied, the retrofit measures must be redesigned or an alternative retrofit strategy with a different PO may be employed. This may lead to an iterative design procedure, until the acceptance criteria for the selected PO are satisfied. The acceptance criteria for the different damping devices can be estimated based on the previous step-by-step procedures. Concerning the acceptance criteria for building components, for deformation-controlled actions (e.g., moment in beams, moment and shear in walls, axial force in braces), the expected deformation capacity (allowable inelastic deformation limits) of all components should not be less than maximum deformation demands calculated by the analysis. Concerning force-controlled actions (e.g., shear in beams, axial force in columns), the lower bound strengths (capacity) of all components should not be less than the maximum analysis forces (ASCE, 2013). Given that there are many conditional details about various types of building components (e.g., steel components and concrete components), the user is encouraged to directly refer to tables presented in the standard ASCE 41-13 (ASCE, 2013) for the different structural components. Note that it is important to understand the economic implication when checking the feasibility of the retrofit design. In case the retrofit strategy is not economical, there are three choices possible (ASCE, 2013): • • •

Considering more refinements in analysis Designing a different retrofit scheme Considering a different PO

ASCE (2013) defines different acceptance criteria depending on if the component is deformation (ductile) or force controlled (nonductile). The relative acceptance criteria were defined by looking at the force
deformation curve as shown in Fig. 5.42 (Step 5, Section 5.5.3.5). Component capacities are defined as given in Table 5.33, while the deformation capacity limits are defined for the structural materials in ASCE (2013).

5.5.3.10 Step 10. Quality control, maintenance, and inspection requirements After the retrofit design is completed, the construction documents including the requirements for construction quality assurance can be prepared. The quality of construction should be checked by the designer who is responsible for the seismic retrofit of a building. For this reason, the plan and requirements of the construction quality assurance are recommended in ASCE 41-13 (ASCE, 2013). The quality assurance plan (QAP) may recognize the work components which should be considered for quality assurance procedures; moreover, it may identify special inspection, testing, and observation requirements to confirm construction quality. The QAP may contain a plan to modify the retrofit design

421

422

CHAPTER 5 Design procedures for tall buildings

to take into consideration unexpected conditions found during the construction phase. Reader can refer to Section 1.5.10.1 of ASCE 41-13 (ASCE, 2013) for more details. In addition to the preparation of QAP, the designer will be responsible for conducting periodic structural observations of the retrofit work process at significant stages of construction (i.e., to visually observe the work in accordance with the construction documents and also to confirm the conditions considered during design). Moreover, an additional investigation, testing process, and reporting should be performed under responsibility of a special inspector (ASCE, 2013). For more details about testing requirements, see the different step-by-step procedures as previously revised as well as in Chapter 7.

5.5.4 CASE STUDY: RETROFITTING EXAMPLES OF HIGH-RISE BUILDING WITH DAMPING SYSTEMS 5.5.4.1 Retrofit strategies for a 35-story building in San Francisco with viscous dampers In 2015 the Pacific Earthquake Engineering Research Center published a project (Lai et al., 2015) on the retrofit of a 35-story steel-framed building. A two-stage retrofit strategy was explored to eliminate deficiencies, identified based on the Tier 3 method of ASCE 41-13 (ASCE, 2013), as follows: •

•

Retrofit stage 1. Replacing the heavy exterior cladding with a lightweight curtain wall system; retrofitting the column splices; and retrofitting the beamto-column connections Retrofit stage 2. Incorporating fluid viscous dampers utilizing a simplified optimization method to optimize damper properties, damper distributions, and damper placement configurations

Retrofit stage 1 had been demonstrated (Lai et al., 2015) to be sufficient for serviceability under BSE-1E hazard-level events, but was not insufficient under BSE2E events. Therefore, the retrofit stage 2 was selected as the most optimal solution. The authors selected an effective damping based on a spectrum modification method (developed by Rezaeian et al. (2012)) that lead to a total damping ratio of 10% and 15% for the X- and Y-direction, respectively, for the BSE-2E hazard level. To provide this level of supplemental damping, damper constants (for both linear and nonlinear ones) were determined based on strain energy method (see Chapter 4). For the nonlinear dampers, the velocity exponent 0.35 was first used, then further optimized based on a parametric study. The initial location of viscous dampers was based on storywise placement and three configurations were selected (Fig. 5.44). Based on preliminary analysis the authors (Lai et al., 2015) the building using uniform distribution (Scheme I) of dampers, but large devices were required at upper stories to large forces near the building base. Scheme II
retrofit resulted in damper dependent upon the ground motions analyzed. Scheme III (i.e., damping proportional to story stiffness) was

5.5 Retrofit of Existing Buildings

FIGURE 5.44 Three distribution schemes for damping constant along building height. Adapted from Lai, J.-W., Wang, S., Schoettler, M.J., Mahin, S.A., 2015. Seismic Evaluation and Retrofit of Existing Tall Buildings in California: Case Study of a 35-Story Steel Moment-Resisting Frame Building in San Francisco. PEER Report No. 2015/14. University of California, Berkeley.

the most effective in reducing story displacement, drift, peak floor acceleration, and beam-to-column connection failure. Additional consideration in damper placement was obtained from architectural constraints, since exterior frames are more suitable than interior ones that are usually adjacent to stairs and elevator boxes (Lai et al., 2015) (see also Section 5.2.1.6 for further considerations). Based on this preliminary study further refinement was conducted. Therefore, additional parametric studies demonstrated that lower stories were more effective for placing dampers compared to upper stories, due to larger deformations at these levels, as shown in Fig. 5.45. These refinements lead to a final total number of dampers of 172 (from a starting estimation of 272). Further optimization studies were conducted based on the reduced configuration of Fig. 5.46. This was based on the combination of PI and cost-related index (CI). The first one was defined as the linear combination of interstory drift ratio,

FIGURE 5.45 Plan and exterior frames equipped with viscous dampers. Adapted from Lai, J.-W., Wang, S., Schoettler, M.J., Mahin, S.A., 2015. Seismic Evaluation and Retrofit of Existing Tall Buildings in California: Case Study of a 35-Story Steel Moment-Resisting Frame Building in San Francisco. PEER Report No. 2015/14. University of California, Berkeley.

423

424

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.46 Elevation of two damper distribution cases: (A) distributed in multiple bays and (B) distributed in corner bays. Adapted from Lai, J.-W., Wang, S., Schoettler, M.J., Mahin, S.A., 2015. Seismic Evaluation and Retrofit of Existing Tall Buildings in California: Case Study of a 35-Story Steel Moment-Resisting Frame Building in San Francisco. PEER Report No. 2015/14. University of California, Berkeley.

residual drift ratio, floor acceleration, and peak floor acceleration. Its value ranges from 0 to 1 where higher values means a better performance. Instead, the second index was estimated as a weight combination between the peak damper force and number of dampers. A smaller CI indicates a cheaper damper solution. To combine both effects a new index was defined subtracting PI and CI. This was identified per each story and the larger the index the more effective the dampers were. Additional parametric investigations were conducted, as follows: •

•

Nonlinear viscous damper velocity exponent. A study was conducted with four values of α (0.2, 0.5, 0.8, and 1) for a fixed damping constant and Scheme III distribution (Fig. 5.45). The results demonstrated that using α 5 1 (i.e., linear damper) is most efficient in reducing floor accelerations, while smaller α is more efficient in controlling story drift ratios. Furthermore, by investigating the effectiveness indices (i.e., PI and CI), it was reported that α in the range of 0.2
0.5 was more desirable in retrofit of such a high-rise building using fluid viscous dampers. Bracing stiffness. The bracings should be stiff enough to ensure that viscous dampers are out of phase with the structural response (Lai et al., 2015). A parametric study was adopted to select an optimal bracing stiffness based on a simplified SDOF model with a certain supplemental damping ratio. Several values of effective bracing stiffness (compared to story stiffness) were assessed, and a reasonable range of stiffness was chosen. Verification of the assumption was conducted through NLTHAs. It was reported (Lai et al., 2015) that where supplemental damping ratio was between 10% and 20%, the choice of a bracing stiffness equal to twice the story stiffness gives acceptable results.

5.5 Retrofit of Existing Buildings

•

Damper horizontal location. In order to ascertain the efficiency of dampers with respect to the plan location, a comparative study was carried out. To this end, two damper distribution cases were selected: distribution in multiple bays and distribution in single bay (see Fig. 5.46). Comparing peak story drift ratios obtained from two damper distributions (Lai et al., 2015) showed that the multiple-based distribution is more efficient in controlling the structural response but produces large forces. Note that concentrating damper along corner bays (Fig. 5.46B) led to larger axial forces (due to accumulation of damper forces) in columns adjacent to dampers. Therefore, the first damper distribution (Fig. 5.46A) was found more efficient.

5.5.4.2 Retrofit strategies for a 24-story building in Osaka (Japan) with oil dampers A 24-story building located in Osaka station was retrofitted using oil dampers. There are 26-m-long columns around the atrium and 10-m-long columns at sixth floor. For these reasons, the lateral stiffness of lower six stories was less than that of upper stories. Hence, it was decided to incorporate supplementary oil dampers in order to amend such a deficiency. A total of 48 oil dampers were adopted in the building with Chevron configuration spatially distributed. This type of damper was chosen based on the following reasons (Tanaka et al., 2003): • • •

No residual (permanent) displacement after damper functionality. Reduced required space for damper. Therefore, the use of displacementdependent dampers was not appropriate in this case. Smaller variations in initial (elastic) stiffness and natural period of building in comparison with metallic dampers.

The building was modeled with a lumped-mass scheme in which for the oil dampers the classical Maxwell model (a spring and a dashpot coupled in series, Section 5.2.1.8.1) was employed. Based on this model, for the design of oil dampers, first the total stiffness of supporting component (spring) was determined. Then, the damping constant and the peak force that can be sustained by damper was specified. The influence of variation in damper properties was also considered in the design process. For this aim, the minimum and maximum of (nominal) properties accounted for error in production and atmospheric temperature were considered, totally resulting in variations with 212% and 110%, respectively. Based on the time-history analysis, with several ground motion histories, the drift ratio acceptance criteria was met.

5.5.4.3 Retrofit strategies for a 30-story building in Osaka (Japan) with viscous dampers A 120-m high-rise building located in Osaka in Japan (Fig. 5.47), which is adjacent to an extension building of 70 m height, was retrofitted using viscous dampers connecting the main building to the extension one. The main structural

425

426

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.47 Perspective of main and extension building (Matsumoto et al., 2012).

system is steel frame. The choice of such a connection is due to different dynamic characteristics (mass, stiffness, and natural period) of the two buildings. For the design of connecting viscous dampers, two investigations were conducted. In the first one, the influence of the damper floor location on seismic responses was studied. Hence, three models with different damper floor locations were investigated (Fig. 5.48A). In the second one, three models were considered (Fig. 5.49B) to analyze the influence of degree of distribution of dampers (with the same damping constant). The buildings were modeled using lumped-mass systems and connecting (linear) dampers by the Maxwell model (Section 5.2.1.8.1). A series of parametric analyses were conducted with variable damping constant and based on two seismic inputs in accordance with the Japanese code design spectrum. The results are shown in Fig. 5.49 as a function of the energy dissipation ratio (the ratio between energy dissipation of dampers and input energy due to seismic waves) for both cases. In the figure it can be seen that there is a local maximum ratio corresponding to an optimum damping coefficient. Moreover, locating dampers at upper floors dissipated more energy.

FIGURE 5.48 (A) Investigation 1 and (B) investigation 2. Adapted from Matsumoto, T., Akita, S., Amasaki, T., 2012. Seismic performance improvement of the existing high-rise building by connecting to its high-rise extension using viscous dampers. In: Proceedings of the 15th World Conference on Earthquake Engineering, Lisbon, Portugal.

FIGURE 5.49 Enerdy-dissipation ratio versus damping constant for models of (A) Investigation 1 and (B) investigation 2 (Matsumoto et al., 2012).

428

CHAPTER 5 Design procedures for tall buildings

FIGURE 5.50 (A) Plan view of installed connecting dampers and zoomed details and (B) Elevation view of frames connected by dampers. Adapted from Matsumoto, T., Akita, S., Amasaki, T., 2012. Seismic performance improvement of the existing high-rise building by connecting to its high-rise extension using viscous dampers. In: Proceedings of the 15th World Conference on Earthquake Engineering, Lisbon, Portugal.

Based on these studies, 12 dampers were installed at 14th and 15th floors (Fig. 5.50). The dampers were positioned at 45 degrees with respect to the principal axes of building in order to ensure identical performance in all directions.

5.5.4.4 Retrofit strategies for a 27-story building in Osaka (Japan) tuned mass damper A 130-m high-rise building with an isolating floor at third floor was retrofitted with two TMDs located at the top due to occupant comfort problems. The designed TMDs included: •

Ice thermal storage tank as the moving mass. The mass ratio obtained was 4.2% in Y-direction and 2.2% in torsional direction.

5.6 Dynamic Modification Devices Strategy Optimization

• • •

Flat steel bars as suspension system (to enable movement of mass from elastic deformation of bars and to get a frictionless system). Rotational slide bearing (to connect flat bars to frame and moving mass). Oil damper (which has the force proportional to square of stroke velocity in order to prevent excessive displacements). The designed TMDs were tested (on-site) as follows:

• •

Static loading test (by pushing the moving mass by oil jacks and determining the force
displacement loop). Free vibration test (by initially displacing one of TMDs and determining building top acceleration and TMDs displacement).

Using TMDs dynamic characteristics identified based upon on-site tests, structural analyses under strong winds were performed. Accordingly, the habitability PI (maximum horizontal acceleration) was determined against the frequency, leading to good performance (response reduction) in both Y- and torsional directions.

5.6 DYNAMIC MODIFICATION DEVICES STRATEGY OPTIMIZATION 5.6.1 INTRODUCTION The effectiveness of dynamic modification systems may differ depending on various parameters such as the type of damping system, damper configuration, damper distribution (location) in height/plan, and damper parameters. Concerning the selection of a suitable type of dampers, as well as the factors affecting how (configuration) and where (location) dynamic modification devices can be placed in buildings, some guidelines are presented in Section 5.2.1.6 for distributed-type dampers (viscous, viscoelastic, friction, and hysteretic), in Section 5.2.2.7 for mass-type dampers (TMD, TLD, and TLCD), and in Section 5.3.1.5.2 for baseisolation systems. Moreover, the dynamic modification parameters can be determined based upon initial requirements (e.g., supplemental damping ratio) that are selected by the designer; some iterations may be used for further refinement of structural response and of damper characteristics. Despite what already discussed in the previous sections and chapters, the dynamic modification device distribution (location) and dynamic modification device parameters may be further optimized with the help of more sophisticated optimization strategies such as various types of iterative algorithms or simpler (noniterative) means. Hence, in Section 5.6.2, the optimization algorithms available in literature, suitable for different kinds of dynamic modification systems, are addressed. Moreover, some more simplified strategies (Section 5.6.3) that are not based on complicated algorithms are briefly reviewed.

429

430

CHAPTER 5 Design procedures for tall buildings

5.6.2 ALGORITHM-BASED OPTIMIZATION PROCEDURES An extensive number of optimization methods based on algorithms have been presented in the literature to identify optimal parameter and placement of dynamic modification devices. In general, the most common ones can be categorized as gradient-based algorithms; GAs; control theory
based algorithms; and heuristic algorithms. The major applications of these optimization algorithms have been related to distributed-type damping systems, especially viscous dampers. In the following, the different procedures are briefly reviewed. Interested readers should refer to the references provided.

5.6.2.1 Gradient-based algorithms In order to determine the optimal device placement, a gradient-based optimization procedure based upon the optimality criteria and relevant performance sensitivities can be used (Takewaki et al., 2012). For a better understanding, various steps of this procedure are illustrated schematically in Fig. 5.51. In such a procedure, sensitivity analyses of the objective function should be usually performed with respect to the dampers’ damping coefficient to find the highest performance sensitivity. This method is subject to iterations (i.e., repeating the sensitivity analyses and finding the highest performance sensitivity) until the required total amount of supplemental damping is obtained. This category of algorithms usually requires programming. Typical gradient-based algorithms are often relative to linear structural behavior. Several studies have been conducted in the past for different device categories, such as: •

•

• • •

Viscous dampers (e.g., Balling and Pister, 1983; Takewaki, 1997, 1999, 2000; Singh and Moreschi, 2001; Mahendra and Moreschi, 2001; Uetani et al., 2003; Lee et al., 2004; Lavan and Levy, 2005, 2006; Attard, 2007; Aydin et al., 2007; Cimellaro, 2007; Viola and Guidi, 2009; Aydin, 2012; Adachi et al., 2013; Lavan, 2015) Viscoelastic dampers (Hwang et al., 1995; Singh and Moreschi, 2001; Mahendra and Moreschi, 2001; Lee et al., 2004; Park et al., 2004; Fujita et al., 2010) Hysteretic dampers (e.g., Uetani et al., 2003) TMDs (e.g., Zuo and Nyfeh, 2004; Wang et al., 2009; Salvi and Rizzi, 2014) TLCDs (Taflanidis et al., 2007)

5.6.2.2 Genetic algorithms GAs are usually efficient to identify the optimal location of dynamic modification devices in buildings and they do not restrict to have a linear behaving structure. Such methods are found more suitable for problems where the PI is not a continuous function of the design variables (e.g., damping constant) (Singh and Moreschi, 2002). One of the advantages of this approach is the possibility to set a

5.6 Dynamic Modification Devices Strategy Optimization

FIGURE 5.51 Representative schematic diagram of gradient-based optimization procedures. Adapted from Takewaki, I., 2009. Building Control with Passive Dampers: Optimal Performance-Based Design for Earthquakes. John Wiley & Sons (Asia), Singapore.

given dynamic modification device capacity, usually based on commercially available solutions. Consequently, the optimal location of a certain number of dynamic modification devices, with the fixed capacity selected, can be identified (Singh and Moreschi, 2002). As a shortcoming, these methods usually require long computational time demand (Singh and Moreschi, 2002). The GA-based optimization techniques have been employed for the identification of optimal placement of various dynamic modification devices, such as: •

Viscous dampers (e.g., Furuya et al., 1998; Singh and Moreschi, 2002; Wongprasert and Symans, 2004; Tan et al., 2005; Dargush and Sant, 2005; Silvestri and Trombetti, 2007; Lavan and Dargush, 2009; Kargahi and

431

432

CHAPTER 5 Design procedures for tall buildings

•

• • • • •

• •

Ekwueme, 2009; Shin, 2010; Apostolakis and Dargush, 2010; Hejazi et al., 2013; Greco et al., 2016) Viscoelastic dampers (e.g., Singh and Moreschi, 2002; Park and Koh, 2004; Dargush and Sant, 2005; Movaffaghi and Friberg, 2006; Shin, 2010; Qu and Li, 2012) Metallic dampers (Dargush and Sant, 2005; Shin, 2010) Hysteretic dampers (Moreschi and Singh, 2003; Ok et al., 2008) Friction dampers (Moreschi and Singh, 2003; Miguel et al., 2014) BRBs (Farhat et al., 2009) TMDs (e.g., Hadi and Arfiadi, 1998; Arfiadi and Hadi, 2011; Singh et al., 2002; Marano et al., 2010; Mohebbi and Joghataie, 2011; Fu et al., 2011; Huo et al., 2013; Herve´ Poh’sie´ et al., 2015; Venanzi, 2015; Greco et al., 2016) Tuned liquid dampers (Ahadi et al., 2012; Chakraborty and Debbarma, 2016) Isolation systems (Pourzeynali and Zarif, 2008; Charmpis et al., 2012; Xu et al., 2013)

5.6.2.3 Control theory
based algorithms Control theory
based algorithms are advantageous due to their capability in reducing the computational efforts because they do not require to compute the structural response using dynamic analyses. The simplified sequential search algorithm (Lopez-Garcia, 2001; Garcia and Soong, 2002) and analysis-redesign procedure (Levy and Lavan, 2006) are among the simplest examples of this kind of algorithms. As noted by Whittle et al. (2012), a disadvantage of such methods may be the lack of PBD criteria within the methods. The control theory
based algorithms have been frequently used for viscous dampers (Zhang and Soong, 1992; Gluck et al., 1996; Lopez-Garcia, 2001; Yang et al., 2002a,b; Ribakov and Reinhorn, 2003; Main and Krenk, 2005; Levy and Lavan, 2006; Cimellaro and Retamales, 2007; Aguirre et al., 2013) and viscoelastic dampers (Zhang and Soong, 1992; Loh et al., 2000; Lopez-Garcia, 2001).

5.6.2.4 Heuristic algorithms In addition to aforementioned algorithms, some other heuristic ones are applied in the literature for the purpose of optimizing dynamic modification systems, such as: •

• • •

Viscous dampers: performance-based heuristic approach (Liu et al., 2005) and artificial bee colony algorithm combined with a gradient-based algorithm (Snomez et al., 2013) Hysteretic dampers: adaptive smoothing algorithm (Murakami et al., 2013). Friction dampers: backtracking search optimization algorithm (Miguel et al., 2015). TMDs: bionic algorithm (Steinbuch, 2011), particle swarm optimization (Leung and Zhang, 2009), harmony search method (Bekdas and Nigdeli, 2011, 2013), ant colony optimization (Farshidianfar and Soheili, 2013), evolutionary operation (Islam and Ahsan, 2012), and charged system search (Kaveh et al., 2015)

5.6 Dynamic Modification Devices Strategy Optimization

5.6.3 NONALGORITHM-BASED OPTIMIZATION PROCEDURES Generally speaking, optimization techniques of dynamic modification systems based on algorithms can be time-consuming as well as complicated for structural engineers. Hence, simpler strategies could be alternatively utilized. Some possible strategies presented in the literature are listed below and are briefly explained: •

Optimality with maximization of supplemental damping ratio (viscous and hysteretic dampers). To locate viscous dampers at more suitable locations, several authors suggested to incorporate them such that the modes supplemental damping ratio, especially the fundamental one, are maximized (i.e., for the same damper properties the higher value of supplemental damping ratio is achieved). According to Ashour and Hanson (1987), such a maximum supplemental damping ratio was achieved once dampers were located in the first story, provided that the building is a shear-type one (e.g., concrete moment frame). Furthermore, to find the optimal placement, Liu et al. (2004) recommended to investigate the influence of various damper configurations on different PIs (e.g., story drift and cost of device) under earthquake. Hahn and Sathiavageeswaran (1992) suggested to place dampers within the lower half floors of shear buildings if story stiffnesses are uniform. For what regards hysteretic dampers in steel moment frames, Inoue and Kuwahara (1998) developed the optimal condition for the device strength and stiffness. The optimization was based on finding the optimal ratio (β, of damper shear strength to frame maximum resistance of the frame), to maximize equivalent viscous damping. This optimal ratio is a function of relative stiffness (k) between the damper and frame. Fig. 5.52 shows the

FIGURE 5.52 Optimum damper’s strength ratio (β opt ) and optimum trigger level coefficient (ψopt ) with respect to relative stiffness (k). Adapted from Inoue, K., Kuwahara, S., 1998. Optimum strength ratio of hysteretic damper. Earthquake Eng. Struct. Dyn. 27 (6), 577
588.

433

434

CHAPTER 5 Design procedures for tall buildings

•

•

•

•

proposed optimal curves for β and the trigger level coefficient (ψ, strength of the entire system (frame plus dampers) at initial yielding). Interstory drift proportional distribution (viscous dampers). Tsuji and Nakamura (1996) proposed some design guidelines for optimal placement of viscous dampers in shear buildings. They recommended to first place the dampers in half the stories in which interstory drifts are greater. Then, according to such a placement, the damping coefficient of dampers can be determined in proportion to the distribution shape of story drifts. Quasioptimal distribution. It is possible to first optimize the stiffness distribution along the building height; then, the damping distribution could be directly determined by considering that damping is proportional to stiffness (Connor and Klink, 1996; Connor et al., 1997). Optimal placement with controllability index (viscoelastic dampers). Shukla and Datta (1999) employed a controllability index (based on the root-meansquare value of interstory drifts) to find optimal locations of viscoelastic dampers in buildings. Accordingly, they stated that the optimal placement depends on the excitation process (e.g., narrowband or broadband excitations) and the modeling type of dampers (e.g., Kelvin model or Maxwell model). Mass-proportional damping (MPD) distribution. Viscous dampers can be placed in shear-type buildings by considering the mass proportionality of damping (see Fig. 5.53) (Trombetti and Silvestri, 2004, 2006; Silvestri and Trombetti, 2007). It was demonstrated by Trombetti and Silvetri (2004) that MPD provides higher value of the first modal damping ratio among Rayleightype damping systems and leads to more optimized solution when dealing with seismic design. The main disadvantage of such a damper distribution is

FIGURE 5.53 Various types of mass proportional damping systems. Adapted from Trombetti, T., Silvestri, S., 2004. Added viscous dampers in shear type structures: the effectiveness of mass proportional damping. J. Earthquake Eng. 8, 275
313.

5.6 Dynamic Modification Devices Strategy Optimization

•

that fixed points are required to install devices and this may not be always attainable in practice. The application of the MPD in a case study 18-story concrete-core and steel-frame building located in Italy is illustrated by Trombetti and Silvestri (2004). They proposed two solutions: 1. Using long BRBs to connect each story to the ground as shown in Fig. 5.53A. Such long systems may be constructed using unbonded braces (Clark et al., 1999a,b,c) or long prestressed steel cables coupled with silicon dampers. 2. Locating dampers between the structure and a very stiff vertical element adjacent to or inside the building; see Fig. 5.53B. Elevator/stairs concrete cores are a typical instance of such an element (Trombetti and Silvestri, 2004). Noniterative optimal design for brace-damper system. London˜o et al. (2012) developed a simple (noniterative) procedure for optimal design of viscous dampers (Maxwell-type model, Section 5.2.1.8.1) when installed in an SDOF system. They considered two criteria for sizing the damper and brace: optimum damper size (ODS) criterion and optimum brace stiffness (OBS) criterion. Having selected a target supplemental damping ratio (ζ s ) due to dampers, two controlling parameters (α 5 ratio between damper stiffness and SDOF stiffness and β 5 ratio between damping coefficient of damper and of SDOF system, multiplied by inherent damping ratio) can be determined using Fig. 5.54. The points associated with ODS and OBS are shown on the plot.

FIGURE 5.54 Added damping ratio map for a structure with 5% inherent damping (London˜o et al., 2012).

435

436

CHAPTER 5 Design procedures for tall buildings

•

•

•

Then, the damper and brace optimum size can be simply determined based on the values of α and β and given SDOF system parameters. Castaldo and De Iuliis (2014) proposed an optimal integrated cost-effective seismic design for viscous dampers supported by braces (Maxwell-type model, Section 5.2.1.8.1) in SDOF systems. The optimized objective function to be minimized was the cost index composed of the cost of SDOF system stiffness and of damper stiffness and damping. Optimality of TLCD parameters. Gao and Kwok (1997) stated that using larger ratio between vertical and horizontal cross-sectional areas in TLCDs can significantly decrease the length of damper; this is a benefit for the application of such devices in flexible structures. Explicit expressions were proposed for optimal parameters of TLCD (damping ratio, frequency ratio, and head loss coefficient) when the structure is undamped (Yalla and Kareem, 2000; Shum, 2015). For the damped case, numerical solutions were proposed by Yalla and Kareem (2000) and closed-form relations by Shum (2015). It was stated that for systems with light damping, the optimal damping ratio of TLCD does not depend on the inherent damping of main system under purely white-noise excitation (Yalla and Kareem, 2000). Ghosh and Basu (2007) developed an explicit relation to estimate the optimum tuning ratio of TLCDs, applicable for lightly to moderately damped structures under base earthquake excitations. The optimization was based upon the minimization of the peak transfer function of the displacement response of the structure. Optimality of TMD tuning ratio. Ghosh and Basu (2007) proposed a closedform expression to determine the optimal tuning ratio (frequency ratio) of TMD, when the main structure is damped. The base of this optimization was to minimize the peak displacement transfer function of the structure. Optimality of midstory isolation systems’ parameters. Zhou et al. (2016) developed an approach with the help of a simple two-DOF system as the equivalence of midstory isolated buildings. The analytical relations to estimate the optimal parameters and location of isolation system were derived based on the minimization of peak base-shear response under base excitations.

CHAPTER

Architectural aspects and building system interaction

6

CHAPTER OUTLINE 6.1 Architectural Aspects .......................................................................................438 6.1.1 Distributed Damping Systems..........................................................439 6.1.2 Mass Damping Systems ..................................................................444 6.1.3 Base Isolation Systems ...................................................................447 6.2 Elevators .........................................................................................................450 6.3 Mechanical Systems ........................................................................................455 6.3.1 Basic Considerations ......................................................................455 6.3.2 Mechanical Floors ..........................................................................456 6.4 Fac¸ade ............................................................................................................457 6.4.1 Double-Skin Fac¸ade........................................................................458 6.4.2 Diagrid Fac¸ade ...............................................................................460 6.4.3 Mega Brace Dampers Fac¸ade...........................................................463

Previous chapters have explained the basics of dynamic modification systems for tall buildings starting from the general concepts till understanding the different products available as well as the bases for their design. This process could be clear for most structural engineers, but other professionals do not usually have a flawless understanding on what are the implications of using these devices in tall buildings. This is why this chapter tries to provide a general overview of different aspects that could be relevant to other professionals involved in the design of tall buildings (architects, mechanical, electrical, and plumbing (MEP) consultants, elevator consultants, fac¸ade consultants, etc.). The goal is to provide some general insights on the possible implications and solutions that could be applied to any tall building that utilizes dynamic modification systems as a means for enhancing the building performance as explained in Chapter 3. The discussion presented in the following sections will make clear the importance of coordination between the different disciplines involved in the design. Indeed, building performance targets (as reviewed in Chapter 3) can be more confidently achieved considering the interaction among all the building components (e.g., structural, mechanical, architectural, etc.) (NEHRP, 2015). For this reason, it is important to organize a design team, including all the actors involved in designing and planning a tall building, such as architects, structural engineers, MEP consultants, geotechnical engineers, curtain wall consultants, fac¸ade Damping Technologies for Tall Buildings. DOI: https://doi.org/10.1016/B978-0-12-815963-7.00006-3 Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved.

437

438

CHAPTER 6 Architectural aspects and building system interaction

consultants, elevator consultants, and contractors (with relative subcontractors, specialty trades, manufacturers, and suppliers). The active participation of the building owner is also necessary in order to coordinate the major aspects. The following sections address the various kinds of interactions that may occur between dynamic modification systems and some other aspects of tall buildings (architectural, elevators, mechanical systems, and fac¸ades). Different adopted solutions will be reviewed to give readers a general overview of the possible applications and outcomes in the utilization of the different dynamic modification strategies.

6.1 ARCHITECTURAL ASPECTS The success of architecturally integrated building systems is the result of a comprehensive approach toward actively engaging the different actors under the direction of the architect. The goal is to obtain an appealing result, as well as performing the required service or amenity in total harmony with other building components. Among the research community, this process is known as the three modes of building integration (Bachman, 2003): physical, visual, and performance. From a wider perspective, building integration can be categorized into two typologies: software and hardware. Software refers to the architectural vision that the building needs to convey, and hardware is the means (e.g., materials, building systems) to achieve this. Dynamic modification systems typically fall within the hardware building integration category. The main architectural impact of these devices is the possibility to expose or conceal them in architectural finishes (e.g., fac¸ade, partitions). This decision depends primarily on the owner and his/her perception of the value that it brings to his/her project. For example, dynamic modification devices can be perceived as a marketing advantage since they can bring a positive security perception for the occupants. On the other side, an auxiliary damping device may be perceived as a structural “patch” that is required to make up for a fundamental lack of strength or stiffness in the primary or bare structural frame of the building. The possibility (or not) to display dynamic modification systems depends partly on the occupancy type of the building. The major factors that affect this choice are: • • • • •

Ability of public “audience” to gain access and/or otherwise view it Shape of the building Available space in which the dynamic modification system can be allocated Relative cost related to the dynamic modification system’s position within the building Safety and maintenance considerations

In the following sections, some architectural solutions adopted in tall buildings equipped with dynamic modification systems are reviewed to provide some precedents for future possible applications.

6.1 Architectural Aspects

6.1.1 DISTRIBUTED DAMPING SYSTEMS Distributed damping systems consist of a series of devices (displacement- or velocity-dependent) that are utilized within a building structure, in multiple locations along the building height (see Chapter 4). Some of the major considerations that an architect should take into account are the following: •

•

•

Geometrical configuration: Different configurations for the installation of distributed dampers in tall buildings were illustrated in Chapter 4. Each of them presents some unique features that can be architecturally relevant especially in the case where the device is exposed. Position within the building: Both vertical and horizontal positions of the damping systems are considered since architectural elements can be affected by their location (such as partitions, doors, and windows). Exposed or concealed: Exposed or concealed dampers depend on the effects desired by the architectural expression.

In the following sections, the major adopted solutions that can be found in the literature are briefly reviewed, to provide readers with some examples of the possible architectural precedents that could be utilized.

6.1.1.1 Damper configurations Distributed damping systems adopt different configurations based on the device utilized and the structural solution utilized (major configurations are shown in Chapter 4, Fig. 4.2). In addition, architectural aspects place a major role since these devices span the entire story height creating several constraints in the freedom in the architectural layout. Hence, dampers placement may become difficult, since the positions intended for dampers may be occupied with architectural components (e.g., windows, doors). This constraint could be more stringent especially in the case of building retrofit since the position of the architectural elements is already defined (Hwang, 1998). Despite the great variety of possible geometrical damping configurations, some solutions can give more architectural freedom. For example, viscoelastic coupling dampers (Christopoulos and Montgomery, 2013) can be distributed vertically along shear wall, as shown in Figs. 4.16 and 6.1. Moreover, indirectly connected dampers, such as stud-type and bracket-type systems, offer more freedom to architectural planning due to the lesser space they occupy, as shown in Fig. 6.2 (Kasai and Kibayashi, 2004). Even more freedom is available when beam or columns are disconnected for placing the dampers (Fig. 6.3) (Kanada et al., 2002).

6.1.1.2 Vertical and horizontal distributions Horizontal and vertical distributions of dampers can be informed not only from structural requirements but also from architectural considerations. One example of vertical distribution informed by architectural needs can be seen in the SUT

439

440

CHAPTER 6 Architectural aspects and building system interaction

FIGURE 6.1 (A) Viscoelastic-coupled damper embedded among RC walls, (B) a typical damper view, and (C) elevation of coupled shear walls including dampers (Montgomery and Christopoulos, 2015).

FIGURE 6.2 (A) Stud-type damper and (B) bracket-type damper.

FIGURE 6.3 (A) Column-type damper and (B) beam-type damper.

6.1 Architectural Aspects

FIGURE 6.4 Distributed viscous wall dampers in SUT Building, Shizuoka, Japan. (A) During construction and (B) after completion (Clark et al., 1999b).

Building in Shizuoka, Japan (Fig. 6.4). From the building view under construction (Fig. 6.4A), it can be seen that a checkerboard-like distribution of viscous wall dampers is adopted, leaving the central bay free for placing glass windows (Fig. 6.4B) (Clark et al., 1999b). Architectural requirements might also impact the horizontal distribution of damping devices. For example, placing them on the exterior of a building can be beneficial on the utilization of the interior space but will increase the constraints on the expression of the external fac¸ade (Taylor and Katz, 2007). Therefore the adopted solution needs to be a balance between all these different aspects.

6.1.1.3 Expose or conceal Distributed damping systems are in most cases less architecturally appealing than mass damping systems (see Section 6.1.2); as a consequence, architects can opt to not expose them. However, there are several building examples in which they have been exposed. In this section, a brief overview of some examples in which distributed damping systems are exposed or concealed is given: •

Diagonal and chevron dampers can be installed that are visible from interior rooms (Fig. 6.5) or exposed from the outside of the building (Fig. 6.6). In case dampers are not exposed, they can be concealed within partition walls, for example, drywall partitions, as shown in Fig. 6.7.

441

442

CHAPTER 6 Architectural aspects and building system interaction

FIGURE 6.5 Internally exposed distributed dampers: (A) chevron viscous dampers (PEER, 2015), (B) oil damper in Osaka Bar Association Building (Tokuda and Taga, 2008), (C) diagonal viscous dampers in a 42-story building (Ekwueme et al., 2010), and (D) X-shape friction damper in McConnel Library Galleria (Pall and Pall, 2004).

FIGURE 6.6 Externally exposed distributed dampers: (A) chevron steel damper and (B) bucklingrestrained brace (Charleson, 2008).

6.1 Architectural Aspects

FIGURE 6.7 (A) Diagonal friction damper in drywall partition and (B) chevron friction damper in drywall partition (Balazic et al., 2000).

FIGURE 6.8 (A) Visible brake damper and (B) wall damper in Ark Hills Sengokuyama Mori Tower (Nakai, 2015).

• • •

•

Viscous wall dampers and brake (hysteretic) dampers (Fig. 6.8) were installed in the 47-story Ark Hills Sengokuyama Mori Tower in Tokyo. Buttresses with viscous dampers were used in the 52-story Allianz tower in Milan (Italy) (Fig. 6.9) (see more details in Chapter 8, Section 8.1.14). Viscoelastic dampers between horizontal trusses and exterior steel tube walls (Fig. 6.10) were used in the World Trade Center in New York City (Mahmoodi et al., 1987). Here, the dampers were hidden in the ceiling but might have been exposed to interior view if desired. Damped outrigger systems (Fig. 6.11) are sometimes utilized within technical/ mechanical floors (a building floor where most of the mechanical devices are positioned). In this case, the damping system is hidden from the general public view.

Additional case study examples showing different architectural solutions adopted are reviewed in Chapter 8.

443

444

CHAPTER 6 Architectural aspects and building system interaction

FIGURE 6.9 (A) Overall view of the Allianz Tower indicating two of the buttresses and (B) viscous dampers at the base of a buttress (Mola et al., 2016).

FIGURE 6.10 (A) View of the viscoelastic dampers in the World Trade Center, New York City (Aiken, 2006) and (B) their schematic representation (Mahmoodi et al., 1987).

6.1.2 MASS DAMPING SYSTEMS Mass damping systems consist of a large mass (or a combination of a discrete number of smaller masses) which reduces building motion (see Chapter 4 for more details) through large differential motion (tuned mass damper) or liquid turbulence (tuned liquid damper or tuned column dampers). This large mass poses several architectural complications since it is usually provided in the upper part of a building (to be more effective). Indeed, the top of the building typically has

6.1 Architectural Aspects

FIGURE 6.11 Damped outriggers and building core (A) floor plan and (B) elevation with highlighted mechanical floors. Adapted from Mathias, N., Ranaudo, F., Sarkisian, M., 2016. Mechanical amplification of relative movements in damped outriggers for wind and seismic response mitigation. Int. J. High Rise Build. 5 (1), 51 62.

the best views, and thereby can be one of the most valuable and desirable spaces for the developer. Therefore a mass damping system may compete with expensive penthouse units, observation areas, executive office spaces, and restaurants, as well as with other building systems like mechanical or elevator machine room. Moreover, the tops of today’s tall buildings are becoming more and more diverse and creative in their designs. Often the pinnacle of the building is not just a space with a purpose, be it private or public, but a critical architectural element integral to the esthetic design of the building. More and more building tops are shaped to enhance the wind performance and provide platforms for energy production or support for communication equipment. Some are designed to attain height goals or to complete architectural esthetic compositions, often as mandated by the aspirations of the owner. Therefore the top of every tall building tends to be unique, and some shapes are better suited than others to house dampers. As previously discussed, the damping system can be exposed or concealed based on several factors, such as the damping system type, owner’s choice, or esthetic intent. One of the most appealing solutions that has been adopted in tall buildings is the Pendulum TMD. This can be utilized as an appealing architectural feature as shown in several tall buildings, such as Taipei 101 (Taiwan) and Shanghai Tower (Shanghai), as shown in Fig. 6.12 (see more details in Chapter 8, Sections 8.2.3 and 8.2.14, respectively). In these two examples, to enhance their architectural features, the architects adopted observation decks where the damping system can be viewed as an attraction.

445

446

CHAPTER 6 Architectural aspects and building system interaction

FIGURE 6.12 Mass damper at (A) Taipei 101 (Irwin et al., 2008) and (B) Shanghai Tower (illustration by Sinelab).

Given the architectural aspects discussed earlier, some of the possible architectural considerations for the different available mass damping systems are as follows: •

•

•

TMD pendulum: • Visually most interesting and simplest for the occupant to understand. • Mass can be shaped to add visual attraction. • Relevant height space required even when using a double pendulum approach. • Safety clearance for damper swing and movement. TMD translational/sliding: • Visually less interesting than the previous one, since it has a machine-like appearance. • Mass can be shaped to add visual attraction. • Reduced height space requirement than a pendulum. • Safety clearance for damper swing and movement. • System requires partial or full enclosure for mechanical maintenance (including hydraulic bearings). TLD/TLCD: • Visually the least interesting since liquid is enclosed in sealed concrete tanks. • Less height space requirements than a pendulum. • No clearance requirement for damper swing and movement.

Additional case study examples showing different architectural solutions adopted are reviewed in Chapter 8.

6.1 Architectural Aspects

6.1.3 BASE ISOLATION SYSTEMS Base isolation systems, as defined in Chapter 4, are utilized to uncouple the building motion (above the isolation level) from the external energy input (earthquake). The preeminent architectural concern is the proper detail of the seismic gap at the interface between the base isolation building and the surroundings (Charleson, 2008). Careful consideration should be given to the movement of different (structural and nonstructural) systems in such buildings and their potential interaction due to the expected earthquakes (Mayes et al., 2012). Some of the most interesting architectural-adopted solutions in the past are as follows: •

•

Isolation at the base. In this case, the isolation plane is located beneath the ground floor (Fig. 6.13). A moat is needed in conjunction with an appropriate cover for providing an extensive separation gap. The architect and the engineer must design a flexible service connection (for gas, water, and electricity) crossing the moat. Moreover, the moat cover should be capable of accommodating lateral movement across the gap without providing restraint; see some details in Fig. 6.13A (Charleson, 2008). An additional example is given in Fig. 6.14 for an underground isolation floor including isolators as the combination of rubber bearings and U-shape steel dampers. Isolation above cantilever floor columns. For this kind of isolation system, the isolation plane (set of isolators) is positioned on the first-story columns (Fig. 6.15). The main issue relates to vertical architectural components, such as cladding, crossing the isolation layer that needs to be detailed in order to accommodate lateral movement without inducing any restraint (Charleson, 2008). An example of an isolation plane (exposed) located above columns including horizontal utility services is shown in Fig. 6.15B (Charleson and Allaf, 2012).

FIGURE 6.13 (A) Isolation plane located below the ground floor (Tsuneki et al., 2004) and (B) section of moat and its cover in a seismically isolated building (Charleson, 2008).

447

448

CHAPTER 6 Architectural aspects and building system interaction

FIGURE 6.14 View of the isolation floor based on the ground and rubber bearings in conjunction with Ushape steel dampers (green) (Mayes et al., 2012).

FIGURE 6.15 (A) Isolation plane located above the cantilever ground floor columns (Charleson, 2008) and (B) view of lead rubber bearings from a car parking (Charleson and Allaf, 2012).

•

•

Isolation above first-story walls. The isolators can be located above the firststory walls, as shown in Fig. 6.16 for a tall concrete building. In this example, the isolation plane was exposed (see Fig. 6.16B). Middle-story isolation. The isolation system can be placed at any location along the height of the building, and its position depends on design requirements. Several applications of middle-story isolation have been found in literature, such as: • Shiodome Sumitomo Building (Tokyo, Japan): The isolation layer was considered at an intermediate elevation of the building (Sueoka et al., 2004; Tsuneki et al., 2004). In such a building, the upper structure (portion of the building above the isolation layer) performs as a mass damper for the lower structure (portion of the building below the isolation layer). In addition to isolators, dampers are installed within the isolation layer to absorb excessive vibrations.

6.1 Architectural Aspects

FIGURE 6.16 (A) Overall view of the building and (B) view of isolators at the first floor (Clark et al., 1999b).

•

The noncontinuity between upper and lower structures is an important architectural issue. It can be observed that offices are located in the upper structure (above isolation interface) while atrium and hotel spaces, which require larger spans, are confined to the lower portion. The location of isolation bearings and dampers is chosen in such a way as to not interfere with architectural spaces, egress, and ease of inspection. • First Hills Libadashi building (Tokyo, Japan, Fig. 6.17A): The isolation story (laminated rubber bearing and lead dampers, Fig. 6.17C) isolates the upper part from the lower one to solve vertical irregularity problems (e.g., excessive drift and damage) under seismic excitations (Fig. 6.17B, Zhou, 2001). In addition, the upper structure behaves as an untuned mass damper for seismic response mitigation (Tsuneki et al., 2004). Nakanoshima Festival Tower (Osaka, Japan, Fig. 4.111): The isolation layer and truss system (Fig. 4.111B) are used to separate the building into three parts: lower-level, intermediate-level, and upper-level floors. The seismic isolation layer was used between the hall (located at lower-level floors) and offices (located at intermediate-level floors) to enhance the seismic resistance (Fig. 4.111B). A giant truss including a mega and belt truss is used between intermediate-level office floors and offices located at upper-level floors to transfer forces from upper floors to perimeter columns of the hall (lower-level floors), as shown in Fig. 4.111B. Such a combined structural system increased the degrees of freedom. The plans of upper-level floors, lower-level floors, and isolation floor are shown, respectively, in Figs. 6.18A C. Fig. 6.18A shows that oil dampers are also incorporated among the interior space of upper level floors. These

449

450

CHAPTER 6 Architectural aspects and building system interaction

FIGURE 6.17 First Hills Libadashi building (Tokyo, Japan): (A) overall view of a 14-story building, (B) short-direction elevation, and (C) internal view of the isolation floor (Tsuneki et al., 2004).

•

dampers are employed within exterior parts of the isolation floor in conjunction with high-damping bearings, as shown in Fig. 6.18C. To prevent crushing, a clearance of 65 cm was adopted between the isolated and fixed sides. Swatch Building (Tokyo, Japan, Fig. 6.19): A self-mass damper (SMD) system was developed (Kidokoro, 2004; Arup, 2009). This system is a combination of mass damper (floor plate) and isolation system (combination of high-damping rubber and sliding bearing, as shown in Fig. 6.19B). Contrary to the addition of mass (i.e., conventional TMD), the SMD system utilizes the existing mass of floors and allows the roof space (in addition to floor space) to be used for the architectural program.

6.2 ELEVATORS Elevators (lifts) provide a suitable hoisting system which move people and goods within a building structure, in accordance with the requirements of the vertical

6.2 Elevators

FIGURE 6.18 Nakanoshima Festival Tower: (A) plan of upper structure, (B) plan of lower structure, and (C) plan of isolation floor (Okada and Yoshida, 2014).

transportation (VT) specification of a particular building. Elevators are usually identified in terms of their number, location, rated load, and rated speed (Andrew and Kaczmarczyk, 2011). Elevator installations are inseparable parts of office and residential tall buildings and form a pivotal part of the program for the sustainable development, design, and construction of the modern high-rise-built environment. They are an absolute necessity for aged and disabled persons. Also, the use of elevators for the evacuation of high-rise buildings and structures during natural and humancaused emergencies has become an important issue (ISO/TC 178 N 484, 2006). The performance of elevator systems deployed in the tall-built environment can be substantially affected by large wind- or seismic-induced motions of

451

FIGURE 6.19 Swatch Building (Tokyo, Japan) with isolated SMD system: (A) key structural features, (B) concept, (C) floor plan, (D) rubber bearing, and (E) sliding bearing (Arup, 2009).

6.2 Elevators

building structures. Loads resulting from the building sway excite the elevator system’s flexible components, such as long ropes and belts and cables of variable length, that are installed as a means for car and counterweight suspension and for compensation of tensile forces over the traction sheave. Ropes employed in the elevator governor and safety gear systems, which bring the elevator car safely to rest in the event of emergency and failure of the normal stopping system, are also strongly affected. In addition, lightly tensioned traveling cables used to transmit power and communication signals between the elevator car and control unit may suffer from adverse dynamic behavior. A complex dynamic situation arises when the building is excited near its natural frequency which in turn approaches the natural frequency of the ropes. In this scenario, transient and stationary resonances and modal interaction phenomena take place (Kaczmarczyk, 2012). Large, lateral motions of the ropes occur when coupled with the response of other components such as lift car, counterweight, and compensating sheave assembly. These phenomena severely affect lightly tensioned compensating cables/ropes and traveling cables that sway at the frequency of the building displaying nonplanar responses (ballooning or whirling motions (Kaczmarczyk et al., 2003)). These often swing out and into the elevator and hoistway equipment such as doors, interlocks, switches, rail guide brackets, and compensating sheave assembly, and get entangled (Strakosch and Caporale, 2010). As a result, the elevator designers have to take into consideration the critical characteristics of the building structure before the elevator is produced and installed. These characteristics typically involve the building geometry, the fundamental/first two natural frequencies and the corresponding modal masses, damping ratios, and the story displacements along its height. These data are then used to carry out a frequency analysis in order to predict the critical resonance scenarios that might arise in the system. Relevant dynamic models and simulation techniques are subsequently used to predict the dynamic interaction phenomena that take place during the elevator travel and when the lift car is stationary at various levels in the building. Given all these issues and problematic situations, a suitable passive and/or active control strategy (as discussed in Chapter 4) can then be chosen and applied to minimize the effects of adverse dynamic responses arising in the building elevator system (Kaczmarczyk and Picton, 2013). Thus in order to develop a safe, efficient, and cost-effective high-performance VT technology, the major characteristics concerning the passive, semiactive, and active damping technologies (see Chapter 4) are of vital importance for the elevator engineering practitioners. Generally speaking, clear communication between structural engineers and elevator engineers is critical for the design and performance of VT systems in tall buildings. Practical considerations typical in the coordination between elevators and dynamic modification systems can be described as follows: •

Since elevators are often located within tall building cores, if a mass damper is located on top of the core, then it will increase the height of the building

453

454

CHAPTER 6 Architectural aspects and building system interaction

•

•

•

pinnacle without allowing the elevators to continue up—making the top of the building hard to access. When machine rooms for elevators and dynamic modification system are located at the same floor, for example, roof, the occupied space and position of both shall be determined in agreement with the designer, architect, and elevator engineer. Advances in new technologies in the elevator industry, such as machine-roomless (MRL, Fig. 6.20, Al-Kodmany, 2015) in which the top machine room of an elevator does not exist, may be less subjected to the space competition with the mass damping system. Large lateral displacements during strong earthquakes should be considered for elevator shafts crossing the isolation interface (e.g., Sueoka et al., 2004; GB50011, 2010; NEHRP, 2015). An example is shown in Fig. 6.21A, where the isolation layer is located at middle-story levels. A minimum horizontal clearance (e.g., 50 cm), between the shaft and the elevator, should be adopted in order to accommodate the large displacements of the isolation layer. The appropriate planning between the upper and lower portion of the building allows the elevator shaft to displace horizontally (without interaction) along multiple stories, as illustrated in Fig. 6.21B.

FIGURE 6.20 Gearless, MRL revolution and space saving on the roof (Al-Kodmany, 2015).

6.3 Mechanical Systems

FIGURE 6.21 Middle-story isolation interface and elevator shaft in tall buildings subject to large lateral displacements (Sueoka et al., 2004).

6.3 MECHANICAL SYSTEMS For integrating structural and mechanical systems in buildings, an early determination of the strategies is required in order to retain systems’ consistency and also save time and cost (Charleson, 2008). For this reason, in this section a discussion about the potential interactions between mechanical systems and dynamic modification devices in high-rise buildings is presented.

6.3.1 BASIC CONSIDERATIONS Mechanical and dynamic modification systems’ interaction requirements are barely considered in building codes, and most of the time basic considerations come from standard practice. Some of the most common general recommendations are as follows: •

During the installation or repair of any mechanical system, the dynamic modification system should not be altered or modified. It is recommended to not remove any permanent construction when performing inspection, maintenance, and replacement of mechanical systems (IMC, 2012).

455

456

CHAPTER 6 Architectural aspects and building system interaction

FIGURE 6.22 (A) Flexible piping services passing through isolation layer to better accommodate lateral displacements (Charleson, 2008) and (B) water line with braided section in base-isolated building to allow better movements (FEMA 454, 2006).

•

•

•

•

The restrictions about the installation of nonstructural components (including mechanical systems) at any floor or roof level shall be determined in the design process (NEHRP, 2015). Unwanted interactions between nonstructural components (e.g., mechanical systems) and any other systems (e.g., building elements, dynamic modification elements, or nonstructural systems) placed in their vicinity shall be considered such that no consequential damage (due to failure/contact of the systems) should occur to the fundamental systems (ASCE, 2017a). In order to detect any undesirable interaction, visual inspection of the nonstructural installation may be performed (Gatscher and Bachman, 2012). The dimensions of mechanical components, their locations and relevant attachment points, and effective weight are some of the data that the design professional shall consider (NEHRP, 2015). For these reasons, the space required for the mechanical system shall be carefully planned in conjunction with the dynamic modification system. In seismically isolated buildings, it is important to pay attention to the interaction between the isolation interface and the mechanical systems. Mechanical components that are at or above the isolation interface should resist the maximum lateral forces. Components that cross the isolation interface should be designed for the total maximum displacement and be capable of accommodating large relative displacements due to earthquake, as shown in Fig. 6.22 (e.g., NEHRP, 2009; ASCE, 2017a; NEHRP, 2015).

6.3.2 MECHANICAL FLOORS In tall buildings, it is common practice to have several mechanical floors. The strategy to assign mechanical floors spread along the height is useful to reduce the potential inconsistency in the placement of mechanical equipment. These floors can also be used to locate dynamic modification devices.

6.4 Fac¸ade

FIGURE 6.23 TSD plan adjacent to mechanical room located at the center (Han et al., 2012). Reprinted with permission from STRUCTURE Magazine, June 2016.

In case, mass dampers are designed to be located at a mechanical floor, the location and space requirements of the mechanical room and dynamic modification systems should be decided by the project team: MEP consultants, architects, owner, and structural engineers. Some examples of mass dampers on mechanical roofs are the TLCD system located on the 50th floor in Random House Tower in Manhattan (Tamboli, 2005), the TSD system on the 58th floor of a 60-story tall building in New York (Han et al., 2012), and the TSD system on the 56th floor of the W Downtown Hotel in New York (Fig. 6.23, Han et al., 2012). The dynamic modification system should be positioned in such a way that it does not interfere with the location assigned to other components, for example, mechanical equipment. For instance, if the core is located near the center of the plan, then the damping system cannot be centrally located and a different solution can be found, as agreed between design team members (see the location of damper and core mechanical service in Fig. 6.23). As an alternative to mass dampers, distributed damping technologies can be allocated. One of the most interesting solutions is the damped outrigger system (Ho, 2016; Smith, 2016), as illustrated in Fig. 6.11 and discussed in Chapter 4 (Section 4.1.1). Incorporating damped outriggers across mechanical floors (rooms) requires adequate coordination. In this case, access to mechanical equipment and service means must be considered to prevent any conflict between the outrigger and mechanical room facilities (Choi and Joseph, 2012). In addition, deeper outriggers can be employed that incorporate a double-story height (Ho, 2016).

6.4 FAC¸ADE Owners of tall buildings are frequently interested in building structures with a unique shape and fac¸ade system. In this case, it is required to combine several

457

458

CHAPTER 6 Architectural aspects and building system interaction

expertise such as design, fabrication, and installation to address the design and construction of fac¸ades. Moreover, there are other concerns such as esthetics, performance, engineering, and quality control for longevity that affect fac¸ade implementation in tall buildings. Hence, the communication between structural engineer, fac¸ade consultant, and architect helps to properly address all these issues (Daraphet, 2013). From the nonstructural point of view, it is recommended that damage, due to severe shaking in fac¸ade elements, shall be controlled in order to save life and maintain the functionality of the system (NEHRP, 2009). The use of dynamic modification systems may be effective for this mean. Regardless of the dynamic modification systems utilized, the interaction with the fac¸ade in tall buildings is a relevant subject. In the following sections, the major research developments in this area are reviewed.

6.4.1 DOUBLE-SKIN FAC¸ADE Tall buildings have seen the application of several different fac¸ade systems, such as glass/metal curtain walls, stone panels, and precast concrete panels. Generally speaking, most of these systems are multilayered, but conventionally no considerable gaps exist between the layers. Alternatively, a different system called double-skin fac¸ade (DSF) with a wide gap between the system layers has been recently proposed (Moon, 2005). The main advantage of the DSF is its improved performance as an environmental intermediary to facilitate the flow for ventilation, leading to more energy-efficient solutions (Moon, 2011). Despite this great advantage, the structural capability of DSF, especially at upper floors in tall buildings, is an important concern that could lead to human discomfort. Therefore dynamic modification systems can be introduced within the DSF to control the excessive movement in tall buildings (Moon, 2011). There are several dissipating systems that may be placed within DSFs (Fig. 6.24): low-stiffness connectors, distributed TMDs, and dissipative elements. Low-stiffness connectors can be installed between the outer skin of the DSF and the primary building structure; see Fig. 6.24A. Under wind excitation, the outer skin including its mass sways significantly, and the vibration in the inner fac¸ade skin (primary structure) becomes fundamentally mitigated. However, this strategy gives a matter of vibration control in tall buildings. The main drawback to low-stiffness connectors is the significant vibration of the outer fac¸ade skin (Moon, 2005, 2011). In distributed TMDs, the outer fac¸ade skins are fixed (similar to conventional DSFs), and additional masses are installed resulting in vertically distributed dampers; see Fig. 6.24B. Such systems effectively control the wind-induced vibration in tall buildings. Although vibration of the outer skin is mitigated, the additional mass is a disadvantage. Compared to single (large) TMD located at the top, distributed TMDs save space at the top of the building. The construction of distributed

6.4 Fac¸ade

FIGURE 6.24 Energy dissipating systems in DSFs: (A) low-stiffness connector and (B) distributed TMD (Moon, 2005).

TMDs is simpler than conventional TMDs since they can be assembled as prefabricated DSF units. Moreover, with the use of TMD DFS damping interaction, smaller damper mass ratio is required compared to the conventional system for the same level of supplementary damping ratio (Moon, 2005, 2011, 2016). Alternatively, dissipative elements (e.g., viscous/viscoelastic dampers) can be used between the wall (main structure) and the structural DSF systems (Fig. 6.25). In this situation, three different schemes can be utilized (Passoni et al., 2014): simple DSF (Fig. 6.25A), bidimensional DSF that stiffens also the fac¸ade at roof (Fig. 6.25B), and tridimensional DSF simultaneously installed along both horizontal directions (Fig. 6.25C). The dissipative elements may be designed to be mainly effective under ultimate limit state earthquakes. In addition to the capability of these elements in mitigating seismic-induced vibrations, they are helpful in reducing the DSF dimensions. Moreover, the potential damage may be lumped into these replaceable dissipative elements; thus postearthquake repair costs and building downtime can be considerably decreased (Passoni et al., 2014). Active systems can also be employed for the DSF system to more efficiently mitigate the vibrations (Fu and Zhang, 2016). In this case, the actuator is attached between the two layers of glass skins of the DSF in order to control the outer fac¸ade movements and depths. Fig. 6.26 illustrates the interaction between active control systems and the fac¸ade accompanied with a close-up view of the connection (Fu and Zhang, 2016).

459

460

CHAPTER 6 Architectural aspects and building system interaction

FIGURE 6.25 Dissipative elements connecting wall systems and DSF: (A) simple double skin, (B) bidimensional system, and (C) tridimensional system (Passoni et al., 2014).

FIGURE 6.26 Interaction of active damping system (damper plus actuator) and DSF system (Fu and Zhang, 2016).

6.4.2 DIAGRID FAC¸ADE A common structural system in which both architectural and structural matters are considered could be the combination of an internal core and external frame such as a diagrid. The diagrid structural systems have diagonal elements instead of vertical columns in order to resist both lateral forces and gravity loads (Moon, 2013). Several studies have been conducted to use this structural system to provide alternative solutions for the implementation of distributed, mass, and base isolation damping systems:

6.4 Fac¸ade

•

•

Distributed damping systems can be installed between floors and the external fac¸ade system (horizontal rings of diagrid system, Fig. 6.27) (Lago et al., 2010). In this study, viscoelastic dampers were utilized to connect the two systems. In using such a damped connection, the force and displacement on the structure may be mitigated, and damage to architectural fac¸ades can be reduced (Lago et al., 2010). In addition to viscoelastic dampers, other types of dampers (e.g., viscous) or isolation systems may be also distributed vertically along the height. Apart from an appropriate performance, the viscoelastic dampers are more suitable for taller buildings, since they require a reduced gap between the fac¸ade and the floors (Lago et al., 2010). Elliptical diagrid systems can be employed as both structural and architectural elements in tall buildings (Arup, 2009). An example is the Mode Gakuen Cocoon Tower, a 50-story school skyscraper in Tokyo, Japan (Fig. 6.28A). Three diagrid frames, surrounding an inner core frame (Fig. 6.28B), are rigidly attached to each other at the top and base of this building. Because of such connections, the bending of each diagrid frame causes considerable shear deformation near the middle stories of the interior core frame. This deformation makes it suitable to utilize shear-type (distributed) dampers to efficiently dissipate seismic response (Arup, 2009). Hence, viscous (oil) dampers were horizontally distributed (six per each floor) from the 15th to the 39th floor of the inner core (where most of the shear deformation was happening). Fig. 6.28C shows schematically an elevation part of the inner core with the installed location of the viscous dampers.

FIGURE 6.27 Interaction of fac¸ade and dampers in a complex-shaped building (Lago et al., 2010).

461

462

CHAPTER 6 Architectural aspects and building system interaction

FIGURE 6.28 (A) View of Mode Gakuen Cocoon Tower in Tokyo with diagrid system, (B) three diagrid frames, inner core frame, and (C) viscous damper elevation within the central core (Arup, 2009).

•

•

Mass damping systems may be employed in order to reduce excessive responses in diagrid systems. One example of this case is related to the use of a friction mass damper (a concrete container including water or sand) on the top four stories of a 72-story building. The damper is attached between the building core (using friction pendulum bearings) and the diagrid structural system (using rubber bearings, Fig. 6.29) (Ramadhan, 2014). Diagrid systems have an intrinsic high stiffness that is quite effective in controlling story drifts, but, on the contrary, it increases the seismic forces. In this case, base isolation systems could be an effective solution to enhance the performance of the diagrid system (Arup, 2009). Fig. 6.30 shows the isolation of the diagrid system with a base isolation system in a 20-story, high-rise building in Tokyo. The isolation layer, executed between the first and second basement floors of the building, consists of multilayered laminated rubber bearings with steel and high-damping elastomeric bearings (to efficiently increase the fundamental period and to have horizontal deformation mainly occurring in the isolation layer). In addition, viscous dampers in the perimeter were utilized to increase the damping in the isolation system and decrease its response (Arup, 2009).

6.4 Fac¸ade

FIGURE 6.29 Topmost stories of a 72-story building indicating (A) diagrid fac¸ade and isolated damper (green block) and (B) plan of damper placed between the fac¸ade and core (Ramadhan, 2014).

FIGURE 6.30 The overall perspective of a 20-story base-isolated building in Tokyo with (A) diagrid system and diamond-type double-skin fac¸ade and (B) isolation of diagrid frame at base (Arup, 2009).

6.4.3 MEGA BRACE DAMPERS FAC¸ADE To improve the energy dissipation capacity of dampers (e.g., viscous dampers), it is possible to install them in mega-brace dampers such that they span several floors in tall buildings (as already discussed in Chapter 4, Section 4.1.1). This results in enlarging the relative displacement between the two ends of the dampers (Zhao and Han, 2016).

463

464

CHAPTER 6 Architectural aspects and building system interaction

FIGURE 6.31 The 181 Fremont tower: (A) view of building and mega brace, (B) schematic representation, and (C) close-up view of dampers (blue) in mega brace (Almufti et al., 2016).

The iconic features of this system can be seen as an architectural appeal for the fac¸ades of a tall building (Almufti et al., 2016). This is the case of the 181 Fremont tower located in San Francisco (Fig. 6.31). This solution was adopted instead of a TMD due to cost-benefit reasons (Almufti et al., 2016). The damper mega brace system adds an 8% supplemental damping that helps reduce both wind and seismic excitations. The system is made up of three braces in one (Fig. 6.31B): primary BRB brace and two external secondary BRB braces with viscous dampers at the ends (see Chapter 8, Section 8.1.10 for more details about this building).

CHAPTER

Testing, inspection, and maintenance

7

CHAPTER OUTLINE 7.1 Codes and Standards Development..................................................................466 7.1.1 US Standards...............................................................................466 7.1.2 Japanese Code .............................................................................470 7.1.3 Chinese Code ...............................................................................471 7.1.4 European Standards .....................................................................472 7.2 Preinstallation Tests and Quality Control .........................................................472 7.2.1 Distributed Damping Systems........................................................473 7.2.2 Mass, Active, Semiactive, and Hybrid Damping Systems..................482 7.2.3 Base Isolation Systems .................................................................486 7.2.4 Testing Examples..........................................................................491 7.2.5 Summary .....................................................................................493 7.3 Commissioning and System Tuning..................................................................493 7.3.1 Distributed Damping Systems........................................................494 7.3.2 Mass Damping Systems ................................................................495 7.3.3 Base Isolation Systems .................................................................500 7.4 Fatigue of Dampers ........................................................................................501 7.4.1 Viscous Dampers ..........................................................................502 7.4.2 Viscoelastic Dampers ....................................................................503 7.4.3 Displacement-Dependent Dampers ................................................504 7.4.4 Mass Damping Systems ................................................................504 7.4.5 Base Isolation Systems .................................................................505 7.5 Building Health Monitoring .............................................................................505 7.5.1 Distributed Damping Systems........................................................507 7.5.2 Mass Damping Systems ................................................................507 7.5.3 Active, Semiactive, and Hybrid Damping Systems ...........................507 7.5.4 Base Isolation Systems .................................................................508 7.5.5 Case Study Examples....................................................................508 7.6 Ongoing Maintenance.....................................................................................513 7.6.1 Standard Recommendations ..........................................................515 7.6.2 Distributed Damping Systems........................................................516 7.6.3 Mass Damping Systems ................................................................516

Damping Technologies for Tall Buildings. DOI: https://doi.org/10.1016/B978-0-12-815963-7.00007-5 Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved.

465

466

CHAPTER 7 Testing, inspection, and maintenance

7.7 7.8

7.9

7.10

7.6.4 Active, Semiactive, and Hybrid Damping Systems ...........................517 7.6.5 Base Isolation Systems .................................................................518 Maintenance Cost ..........................................................................................519 Postextreme Event Inspections........................................................................519 7.8.1 Distributed Damping Systems........................................................521 7.8.2 Mass Damping Systems ................................................................522 7.8.3 Base Isolation Systems .................................................................523 Transportation, Installation, and Care of Works................................................524 7.9.1 Transportation of Damping Systems ...............................................524 7.9.2 Storage/Installation of Damping Systems ........................................525 7.9.3 Care of Works of Damping Systems ................................................526 Resilience-Based Earthquake Design Initiative and United States Resiliency Council Rating Systems .............................................526 7.10.1 Resilience-Based Earthquake Design Initiative Rating System ........527 7.10.2 United States Resiliency Council Rating System ...........................529

This chapter focuses on the testing, inspection, and maintenance of dynamic modification systems and relative components to achieve long-term performance objectives and the expected life of the dynamic modification system. Serving this scope, quality control procedures, consisting of preinstallation tests and long-term periodic inspections, are performed. Preinstallation tests are normally required to be performed to verify assumed properties and acceptance criteria used in the analysis and design. In addition, long-term periodic inspection and maintenance program are required to assure the expected life and performance of the damping system and components.

7.1 CODES AND STANDARDS DEVELOPMENT In the following, the general code requirements outlined for testing, inspection, and maintenance of damping systems and devices are reviewed. The focus of review is related to US standards, however, other international codes such as Japanese, Chinese, and European standards are also reviewed.

7.1.1 US STANDARDS Historically, in United States the design and specification of damping device technologies have largely occurred within specific high-performance industries, such as automotive, aviation, and aerospace; where standard design, fabrication, manufacturing, testing, inspection, maintenance, and certification acceptance criteria and practices have been developed. For buildings, US design codes and standards have been researched and implemented based largely on the development of seismic design using supplemental energy dissipation devices to improve seismic performance. Over the past

7.1 Codes and Standards Development

30 years, research, standard practice, and codes have focused on the development of detailed seismic design requirements for both seismically isolated structures, normally referred to as base isolation, and structures with damping systems. Based on the consensus approved FEMA P-1050 (NEHRP, 2015), ASCE 7–16 (ASCE, 2017a) is the most recent building code standard incorporating these requirements for new buildings. These requirements including supporting commentary and references are contained in ASCE 7–16 (ASCE, 2017a), also in Chapter 17, Seismic Design Requirements for Seismically Isolated Structures, and Chapter 18, Seismic Design Requirements for Structures with Damping Systems. Prior versions of the ASCE 7 standard (ASCE 7–05, ASCE 7–10) included these provisions and were adopted by IBC building code (IBC 2009, 2012, 2015). The ASCE 7–16 (ASCE, 2017a) standard is adopted by the IBC 2018 (IBC, 2018) building code. For existing buildings, the ASCE 41 (ASCE, 2013, and ASCE, 2017b) incorporates similar design and analysis provisions while relaxing requirements for the superstructure intended to facilitate the design and implementation process of seismic retrofit and rehabilitation of existing buildings. These requirements including supporting commentary and references are contained in ASCE 41-17 (ASCE, 2017b), along with Chapter 14, Seismic Isolation, and Chapter 15, Design Requirements for Structures With Supplemental Energy Dissipation. The ASCE 41-17 (ASCE, 2017b) standard is adopted by the IBC (2018) building code. The most recent edition of the ASCE 7-16, Chapter 17 (ASCE, 2017a), provisions for seismic-base isolation contain significant modifications with respect to superseded versions, with an intention to facilitate the design and implementation process of seismic isolation, thus promoting the expanded use of the technology. Instead of addressing a specific method of seismic isolation, the standard provides general design requirements applicable to a wide range of seismic isolation systems. Because the design requirements are general, isolation system hardware is required to be tested to confirm the engineering parameters used in the design and verify the overall adequacy of the isolation system. Use of base isolation systems whose adequacy is not proved by testing is prohibited. In general, acceptable systems (a) maintain horizontal and vertical stability when subjected to design displacements; (b) have an inherent restoring force defined as increasing resistance with increasing displacement; (c) do not degrade significantly under repeated cyclic load; and (d) have quantifiable engineering parameters (such as force–deflection characteristics and damping). The requirements of ASCE 7-16, Chapter 18 (ASCE, 2017a), apply to different types of damping systems including displacement-dependent (hysteretic or friction) and velocity-dependent [viscous or viscoelastic (VE)] damping devices, and base isolation. The damping system (DS) is defined separately from the seismic force– resisting system (SFRS), although the two systems may have common elements. As illustrated in Section 18.1 of the commentary, the DS may be external or internal to the structure and may have no shared elements, few shared elements, or all

467

468

CHAPTER 7 Testing, inspection, and maintenance

elements in common with the SFRS. Elements common to DS and SFRS must be designed for a combination of loads of the two systems. When DS and SFRS have no common elements, the damper forces must be collected and transferred to the members of the SFRS. Compliance with these requirements is intended to lead to a building performance comparable to that of a structure with a conventional seismic force–resisting system. However, the same methods can be used to achieve enhanced or better performance than code minimum objectives. A manufacturing quality control plan for production of dynamic modification devices and components shall be established by the design professional. This plan shall include descriptions of the manufacturing processes, inspection procedures, and testing necessary to ensure the quality control of production devices. At a minimum, the quality control testing program shall evaluate the consistency of results from a specified sample of device units. The results of each test shall be verified to fall within an acceptable range of properties described in the project specifications. These limits shall agree with the specification tolerances on nominal design properties for the average properties of all devices of a given type. The quality control procedures adopted to manufacture of prototype and production damping devices should be similar. In the code standard, two different types of testing procedure are required (for both isolation and damping systems): prototype and production qualification tests. The prototype qualification testing is used to verify the force
velocity
displacement relationships and damping properties assumed for the damping device’s nominal design properties used in the analytical modeling and design of the damping system, and appropriateness for the project and use. From the prototype tests, upper and lower bound damping device properties are established consistent with code specified limits. Alternatively, damping device properties may be based on prior prototype tests of devices of similar type and size meeting similar requirements. Normally, during a project design development, assumed damping device nominal properties are specified, and/or provided by one or more device manufacturers for each type of device used in the damping system. The production qualification testing is conducted prior to installation of individual damping device units in a building used to validate the nominal device properties. It is an essential requirement that the fabrication and quality control procedures used in the manufacture of all prototype and production devices shall be identical. These procedures shall meet all project acceptance criteria and be approved by the registered design professional prior to the fabrication of prototype devices.

7.1.1.1 Prototype qualification tests ASCE 7-16 (ASCE, 2017a) provides detailed requirements for prototype tests for both isolation and damping systems, such as data recording, sequence and cycles of testing, testing similar devices, determination of force
velocity
displacement characteristics, maximum and minimum vertical load, and device adequacy. These tests shall be conducted and approved prior to production of devices for construction.

7.1 Codes and Standards Development

Device nominal properties determined from prototype testing shall meet the acceptance criteria established using upper and lower bound properties as reviewed in Chapter 5 (Section 5.1.3.3.3). These criteria shall establish likely variations in the material properties to account for aging and environmental conditions, manufacture variation, heat, rate of loading and other effects as appropriate. Additionally, independent inspection reports shall be provided summarizing the following aspects: • • • • •

Observations of all prototype tests Data from each conducted test Inspection of the tested damping devices after each test Testing remarks Comparison of results with respect to prototype damper bounded acceptance criteria

The report should also include a description of the equipment used in the prototype testing with the information on calibration of the equipment and instruments used.

7.1.1.2 Production qualification tests The requirements for production qualification testing of isolation and damping devices prior to installation are specified by ASCE (2017a). Properties of devices used in construction shall be determined from the production qualification testing that meet the acceptance criteria established using upper and lower bound properties as established in prototype tests and reviewed in Chapter 5 (Section 5.1.3.3.3).

7.1.1.3 Design review The ASCE 7-16 (ASCE, 2017a) code standard includes mandatory requirements for independent design review for both base isolation and damping systems along with corresponding test programs. Design reviews are performed by one or more individuals possessing knowledge of the following items. At least one of the reviewers shall be a registered design professional. Among all the revisions the following are included in ASCE (2017a): • • • • •

Project design criteria including site-specific spectra and ground motion histories. Device selection and relative design parameters. Device test data and relative property modification factors. Prototype and qualification testing program. Preliminary and final design of the seismic force resisting system and base isolation and/or damping systems.

7.1.1.4 Inspection, periodic testing, and maintenance program ASCE (2017a) provides requirements for the inspection, period testing, and maintenance program of damping and base isolation devices. In order to facilitate this,

469

470

CHAPTER 7 Testing, inspection, and maintenance

means of access for inspection, removal and/or replacement of all devices shall be provided. The registered design professional responsible for the design of the structure shall establish an inspection, maintenance, and testing schedule for each type of device to ensure that the devices respond in a dependable manner throughout their design life. The recurrence of these shall be based on the service history of the devices. Prior to occupancy, the registered design professional shall complete a final series of observations of structure and devices to verify that conditions allow free and unhindered displacement to accommodate maximum total design displacements.

7.1.2 JAPANESE CODE According to JSSI Manual (2003) and Notification No. 1446 (MOLTI, 2000) requirements, the performance of passively damped and isolated buildings should be superior to the general buildings without any dynamic modification systems. For damped structures, the property declaration, quality assurance, and maintenance of damper devices are considered in the Japanese standard code (JSSI Manual, 2003). In this standard, it is recommended that the quality of dampers to be assured with the help of all possible measures. For maintenance purposes, JSSI Manual (2003) suggests long-term warranty, as well as postearthquake inspections of dampers and base isolation systems. Regarding the property declaration, damper’s target performance (under both seismic and wind excitations with various intensity levels) and relative limits should be specified (Kibayashi et al., 2004). More details about testing and inspection of distributed-type dampers based on JSSI manual (JSSI Manual, 2003) are illustrated in the literature (Tanaka et al., 2004; Ishikawa et al., 2004; Nakata et al., 2004) and reviewed mainly in Section 7.2.1. In this regard, it is recommended to perform material tests (similar to prototype test in US standards), damper performance tests (similar to qualification/ production test in US standards), and quality control procedures (including acceptance inspections) from both designer and manufacturer. Regarding quality control of dampers the following list is provided (JSSI Manual, 2003; Tanaka et al., 2004): • • • •

Setting of the target performance. A target performance for the check of quality should be set by the designer. Total control. The quality control is based on standards and the designer or design supervisor who is in charge of this. Construction control. This should be done by the constructor or construction supervisor. The choice of manufacturer can be considered as well. Manufacture control. The constructor and manufacturer are responsible in this case. Manufacturing should be based on standards with the following items to be controlled: material acceptance, device dimensions, weld and assemblage, filling viscous material, device coating, and packing.

In addition, to JSSI manual (JSSI Manual, 2003) for base isolation system, advanced quality control specifications and device qualifications are regulated by

7.1 Codes and Standards Development

Notification 1446 (MOLTI, 2000). This requires that each device should have a stable behavior under design loads and this should be verified through testing. Another important aspect of the Japanese regulations is the possibility to use preapproved devices that have gone through an accreditation process, which requires the testing of minimum three devices (similar to prototype testing in US standards, as shown in Section 7.1.1.1). This allows to cut the costs of specifying requirements for testing on a project-based level (Becker and Furukawa, 2010).

7.1.3 CHINESE CODE In China there are three majors standards (GB50011, 2010; MHURD, 2013; MHURD, 2012) that provide requirements for testing, inspection, and maintenance of dynamic modification devices. The major characteristics of each one are the following: •

•

•

GB50011 (2010) recommends determining the durability and design parameters (e.g., stiffness and damping) of damping and isolation systems with the use of testing. Concerning velocity-dependent devices, the following characteristics shall be specified by testing: design displacement, maximum displacement, restoring force characteristics, various temperature conditions, and loading frequencies. For displacement-dependent dampers, the above mentioned parameters can be determined by repeated static loading. For isolation devices during testing procedures, the stress limit indicated in Chapter 5 (Table 5.1) shall be satisfied. As mentioned in GB50011 (2010), the damping and isolation systems and their components should be simply accessible for maintenance aims. JGJ297-2013 (MHURD, 2013) focuses on the design, fabrication, inspection, and maintenance of structure with dynamic modification systems, which adopt in the structures located in Grades 6
8 seismic regions. This code provides the requirements for the seismic design method for structures using supplemental energy dissipation, including main structural analysis, performance-based design of energy dissipation building, and distribution principle of energy dissipation devices. In addition, the requirements of technical characteristics of energy dissipation devices as well as connecting and details of energy dissipation parts are also provided. For the fabrication and inspection part, considerations for the fabrication, inspection procedures, and site acceptance of devices’ quality are given. In addition, a maintenance program is also required to assure the expected life of damping system. JG/T209-2012 (MHURD, 2012) focuses on the performance requirement of different types of dampers, such as viscous fluid damper, metal yield damper, and bucking-restrained brace. The performance requirement includes material quality, mechanical property, and durability. In addition, the consideration for the preinstallation test method and long-term periodic inspection requirement is also provided.

471

472

CHAPTER 7 Testing, inspection, and maintenance

7.1.4 EUROPEAN STANDARDS In European standards different provisions are given for damping and isolation systems as follows: •

•

For distributed damping systems, EN 15129 (CEN, 2009) addresses the general design requirements, specifications on testing/manufacturing, inspection, and maintenance of this category of devices. Moreover, various kinds of tests for displacement-dependent and velocity-dependent devices are described (more details are given in Section 7.2.1). Damping systems should be constructed and installed in such a way that any inspection, maintenance, and replacement are possible (CEN, 2009). Furthermore, it is stated that viscous dampers should be designed and constructed in such a way that no maintenance is required during their expected life under service loads. For base isolation systems, Eurocode 8 (CEN, 2004) states that sufficient space should be considered between the isolation layer and superstructure in order to facilitate the maintenance, inspection, and replacement. Moreover, it is mentioned that any change in the temperature and in device properties should be accounted for during design. Furthermore, in Section 12 of EN 15129 (CEN, 2009), as well as in Part 10 of EN 1337-10 (2003), two different inspection types are addressed (for base isolation): regular inspection and principal inspection. Additionally, ISO 22762-1 (ISO, 2010) provides recommendations for the test methods to determine properties of the rubber material appropriate for manufacturing and determination of the characteristics of elastomeric seismic isolators.

7.2 PREINSTALLATION TESTS AND QUALITY CONTROL Commercially available dynamic modification devices can have standard features or they require a specific design process and experimental calibration. The second option is the most common situation in the case of tall buildings, given the characteristics of these objects and their enhanced performance requirements. Additional damping devices are usually introduced with a specific goal, such as protection of the structural system from earthquakes or high-intensity winds, or mitigation of wind effects in service life conditions (see Chapter 3). In any case, the dampers need to be able to withstand all the actions resulting from design load combinations (see Chapter 5). This means that even devices designed for comfort, which need to be fully effective for very low vibrations, that is, under wind loads corresponding, for example, to a return period of 1 or 10 years, must still be able to accommodate the displacements/velocities associated with ultimate limit state lateral loads due to the maximum expected wind or earthquake loading. These types of dampers require special care in the manufacturing process and

7.2 Preinstallation Tests and Quality Control

peculiar functional tests on their components, which are not always codified in standards. In the following section, preinstallation testing methods for various kinds of damping systems and isolation systems are described as recommended by standard codes (reviewed in the previous section); moreover, additional recommendations are established based on literature review.

7.2.1 DISTRIBUTED DAMPING SYSTEMS The testing recommendations given by standard codes usually emphasize following types of distributed damping systems: velocity- (viscous dampers), displacement-dependent dampers (steel and hysteretic dampers), and VE dampers. In the case of velocity-dependent viscous dampers used for comfort, the design specifications may incorporate further requirements (e.g., the different system of valves under wind and earthquake, and the type of viscous fluid), in addition to those specified by codes, which generally refer only to seismic protection. In any case, it is always advisable to carry out laboratory tests on prototypes, during the industrial design phase, if there has not been a similar physical size or capacity of damper, to validate the appropriateness of the technology to be used in the project and to validate the design assumptions. Following this, qualification/production tests should be conducted on a representative sample size, in order to guarantee that the as-built product provides the required performances. The following parts describe the main aspects and recommendations of preinstallation tests for distributed-type control systems, as outlined by European standard, US standard, Japanese standard, and Chinese standard.

7.2.1.1 Testing based on US standard In accordance with the last updated version of US standard, ASCE (2017a), the required testing for distributed damping devices are: prototype and production test. The first testing program is done before producing the devices for construction, and the second one is required for all produced devices for any given type and size. ASCE (2017a) provides a list of test sequences for distributed-type dampers that shall be conducted on at least two full-size specimens and for each type and size used. In addition, for each device the in situ gravity loading and thermal conditions shall be properly taken into account. For what regards prototype tests the following are required: •

•

Test 1: 2000 continuous fully reversed cycle for the design windstorm at a frequency equal to 1=ðT1 Þ. This test is not required if the damper is not subject to design wind force or the wind force is less than the damper yield force/slip force in the case of hysteretic dampers. Test 2: Sequence of fully reversed (sinusoidal) cycles at frequency equal to 1=ð1:5T1 Þ for the following magnitude: 10 cycles at 0.33 times the maximum

473

474

CHAPTER 7 Testing, inspection, and maintenance

•

•

considered earthquake (MCER) device displacement, five cycles at 0.67 times the MCER device displacement, three cycles at 1.0 times the MCER device displacement. In the last case, if the damper force in this test is less than the MCER damper force from analysis, it should be repeated at a frequency that produces a damper force equal or greater to the MCER damper force from analysis. Test 3: The above tests shall be repeated for at least one damper at a minimum and maximum temperature (around design temperature), in case the damper properties are temperature dependent. Test 4: Testing for displacements less than or equal to the peak damper displacement and frequency ranging in 1/(1.5T1) to 2.5/T1. If the change in the force-deformation characteristics is higher than 15%, the preceding Test 2 should be additionally conducted at frequencies equal to 1/T1 and 2.5/T1.

Instead, for qualification tests, the following test sequences are required (established by the registered design professional): •

Test 5: 100% of the devices shall be tested for three cycles at 0.67 times the MCER device displacement at a frequency equal to 1=ð1:5T1 Þ.

Given that testing items are essentially identical for prototype and qualification tests, a comparison between results obtained from both testing programs will be possible (ASCE, 2017a). Moreover, the measured testing values shall fall in between the specific tolerances as reviewed in Chapter 5 (Section 5.1.3.3.3). Additional testing requirements and adequacy criteria for various types of distributed damping systems are listed in Table 7.1, based on the recommendations given by ASCE (2017a). Alternative testing methods (other than those mentioned above) are also permitted, if the following conditions are met (ASCE, 2017a): • •

• •

The alternative methods are equivalent to the cyclic testing requirements proposed in the standard code. The dependency of damper performance upon the ambient temperature, loading frequency, and temperature increase due to tests is taken into consideration by the alternative testing methods. The alternative testing methods are approved by the structure-registered professional designer. Conducting a full-scale dynamic testing is impossible because of, for example, test machine limitations (to tackle this limitation, properly reduced-scale prototypes can be used).

7.2.1.2 Testing based on the Japanese standard According to the Japanese standard, JSSI Manual (2003), performance and quality control are crucial to appropriately achieve the complete design of passive control technology. For this reason, in general, two testing programs are

7.2 Preinstallation Tests and Quality Control

Table 7.1 Test Requirements for Distributed-Type Dampers (ASCE, 2017a) Damper Type

Test Requirements

Velocity-dependent dampers

No damage (e.g., leakage, yielding, and breakage) for qualification test. For Tests 1
3, at a given frequency and temperature: The difference between the maximum damping force (and minimum damping force) at zero displacement in every cycle and the average maximum force (and minimum force) at zero displacement obtained from all cycles should be less than 15%. The difference between the area under hysteretic loop of a damper in every cycle and the average one obtained from all cycles should be less than 15%. The designer should check the following quantities: • The average maximum and minimum forces at zero displacement. • The average area under the hysteresis loop (determined from loading each damper with three cycles, at a frequency equal to 1=ð1:5T1 Þ and at the displacement equal to MCER displacement of the damper). These values should be limited by the corresponding minimum and maximum values obtained from the multiplication of lambda factors of specification tolerance (λðspec;maxÞ and λðspec;minÞ )a by the nominal properties.

VE dampers

For Tests 1
3, at a given frequency and temperature, the difference between the effective stiffness of a damper in every cycle and the average one obtained from all cycles should be less than 15%. The designer should check that the effective stiffness is limited by the nominal value multiplied by the minimum and maximum lambda factors for specification (λðspec;maxÞ and λðspec;minÞ )a

Displacement-dependent dampers

For Tests 1
3, at a given frequency and temperature:

a

• The difference between the maximum damping force (and minimum damping force) at peak damper displacement in every cycle and the average maximum force (and minimum force) at zero displacement obtained from all cycles should be less than 15%. • The difference between the average maximum (and minimum) forces at zero and peak displacement and target values determined by the designer, as well as between the average area under the hysteresis loop and target value, should be less than 15%.

Additional details about the lambda factors can be found in Chapter 5 (Section 5.1.3.3.3).

475

476

CHAPTER 7 Testing, inspection, and maintenance

addressed: material and performance tests that are similar to prototype and qualification tests, respectively, from the American standards (ASCE, 2017a). In the following, the requirements of this code are explained for viscous dampers, VE dampers, and steel dampers, independently. In general, all dampers in the Japanese market are always tested in advance to guarantee the appropriateness of the materials used (such as durability, flash point, temperature dependency, durability, fire resistance, and fatigue behavior) and these materials are commonly available for purchase. Therefore for a given project typically only basic performance tests are required. However, when a new product wants to enter into the market, it must demonstrate the appropriateness of the material properties through testing.

7.2.1.2.1 Viscous dampers Tests for viscous dampers according to the Japanese standard (JSSI Manual, 2003) can be classified as follows (Tanaka et al., 2004): •

•

Material test. Three materials should be considered: steel, viscous fluid, and filling material. The steel material should be stainless and in accordance with standards. The physical and chemical properties of the viscous fluid and filling material should be tested, for example, with regard to dynamic viscosity and flash point (the lowest temperature at which the viscous liquid will ignite). Performance test. This is based on the determination of force
velocity
displacement relationship, and on the comparison of damping force at the maximum velocity and the supposed design value. To test dampers, three or more cycles of loading may be adopted at a predefined condition (Tanaka et al., 2004). For example, a possible test is the study of the damper shear resistance temperature
dependence in wall or cylinder type dampers. It is also recommended to test the damping dependence to frequency, velocity, amplitude, and temperature. Moreover, special tests to evaluate the durability of these devices are recommended by JSSI Manual (2003). These include aging alternation, fire resistance, water resistance, heat resistance, and weathering resistance.

The designer and manufacturer should decide about the number of devices and testing conditions (e.g., temperature, frequency, amplitude, and velocity) for the quality control program of viscous dampers. Furthermore, three inspection cases are proposed: • • •

Acceptance inspection: at the time of device manufacturing and also after carrying it to the site Inspection under construction: after damper installation using visual checks Completion inspection: hard to carry out since dampers are rarely accessible due to their covering by finishing material

7.2 Preinstallation Tests and Quality Control

7.2.1.2.2 Viscoelastic dampers Tests for VE dampers according to the Japanese standard (JSSI Manual, 2003) can be classified as follows (Ishikawa et al., 2004): • •

Material test. Testing items may include shear and adhesion performance, shear strain, and fire resistance of VE material. Performance test. Test that measures and confirms the dynamic characteristics (force
displacement
velocity) of VE dampers. In this case, the sensitivity of dampers to strain, temperature, and frequency shall be tested to ascertain the device performance. Moreover, durability evaluations (e.g., fire resistance, moisture resistance, and fatigability tests) are recommended. Based on the results obtained from the performance tests, three limit states can be specified by the designer for each VE damper: • End limit state: determined based on two points (the fracture of material and fatigability) • Damage limit state: for which the device can be repaired for further utilization • Service limit state: defined as the damper is usable continuously without any repair requirements

For device quality control, the manufacturer of VE dampers shall conduct the following tests/inspections prior to shipping (JSSI Manual, 2003; Ishikawa et al., 2004): • • • • • •

•

•

Material tests of VE material Performance test with standard specimens, for example, verification of damping and restoring force features Inspection of VE material quantity, for example, by weighting before and after filling the VE material molds prior to fixing the VE material to steel plates Inspecting the device dimensions Appearance inspection, for example, to verify that VE damper material layers are bonded properly with no peeling or flaws Acceptance inspection: under responsibility of contractor, it may include those inspection items listed earlier or some others (e.g., check of type and number of devices/components, check of no peeling, or flaws) Inspection under construction: also under the responsibility of the contractor to confirm accurate mounting and bolt tightening during installation of the dampers to the structure Completion inspection: to be conducted by the designer along with contractor to confirm the configuration as installed following the mounting procedure

7.2.1.2.3 Steel dampers Tests for steel dampers according to the Japanese standard (JSSI Manual, 2003) can be classified as follows (Nakata et al., 2004): •

Material test. This test is mainly used to determine the mechanical characteristics of steel, for example, yield strength, tensile strength, and elongation.

477

478

CHAPTER 7 Testing, inspection, and maintenance

•

Performance test. This test can be conducted in accordance with the hysteretic curve (at the most critical displacement), elastic stiffness, postyield stiffness, and yield strength. The dependency of device performance on displacement velocity (strain speed) and repetitive cycles shall be considered during testing. Furthermore, the accumulated energy absorption performance can be tested under two classifications: low-cycle fatigue curves (i.e., strain vs cycle) and accumulated plastic deformation test. The durability of devices is not necessarily evaluated, if remedies such as coating are used for rust prevention. In the case, that any plastic deformation occurred in a steel component, potential peeling may occur in the coating; thus the recoating of that component should be performed. The fire resistance of steel dampers is important to be evaluated especially for long time evaluations. An appearance inspection and potential change of device may be required after a fire.

According to JSSI Manual (2003), the quality control of steel dampers is often based on material testing, dimensional inspections, and nondestructive inspections (Nakata et al., 2004).

7.2.1.3 Testing based on the Chinese standard According to the Chinese standard, GB50011 (2010), performance tests (similar to qualification tests in the American standard (ASCE, 2017a), see Section 7.1.3) are recommended for distributed dampers. The testing requirements recommended in this standard are as follows: •

•

Testing sampling • Viscous dampers. A sampling inspection (with 100% qualification rate) should be performed by a third party with 20% of devices (but not less than two devices) of the same type and characteristics. • Other dampers. A random inspection (with 100% qualification rate) should be performed with 3% of devices of the same type and specifications. If few devices of the same size and characteristic exist, then 3% of the devices of the same kind (but not less than two devices) can be inspected. Testing acceptance • Viscous and VE dampers. The difference between the main design parameters and their values obtained from tests should not be higher than 15% after repeating 30 cycles at the design velocity and at the fundamental frequency of structure. • Displacement-dependent dampers. The difference between the main design parameters and their values obtained from tests should not be higher than 15% after repeating 30 cycles at the design drift (displacement).

7.2.1.4 Testing based on European standard EN 15129 (CEN, 2009), “Antiseismic devices,” recommends mainly two general testing programs for distributed dampers: type test and factory production control test that are equivalent of prototype and qualification tests of the American

7.2 Preinstallation Tests and Quality Control

standard (ASCE, 2017a). Depending on the type of damping device, the experimental tests and relevant tolerances and acceptability thresholds, defined in EN 15129 (CEN, 2009), can be adopted not only for earthquake protection dampers but also for wind comfort dampers, if applicable. In the case, wind loads are predominant in design, at least one prototype damper (velocity-dependent type) shall be verified using wind load cycle test (refer to Section 7.4.2.7 of CEN (2009) for more details). The major test requirements, per EN 15129 (CEN, 2009), are the following: •

•

Displacement-dependent devices. The mechanical characteristics of the whole device and its construction material can be verified using type test and factory production control test. For both testing programs, the evaluation of force versus displacement cycle and ramp test is recommended at least for one device. In this case, the governing testing details are given in Sections 6.3.3 and 6.3.4 of EN 15129 (CEN, 2009). Furthermore, the corresponding testing items are summarized in Table 17 of EN 15129 (CEN, 2009). Regarding the material, testing requirements of steel and shape-memory alloys are particularly addressed in Sections 6.3.1 and 6.3.2 of EN 15129 (CEN, 2009). Velocity-dependent devices. Both type test and factory production control test can be utilized. For type tests, the testing items recommended include pressure, low velocity, constitutive law, damping efficiency, wind load cycle, seal wear, and stroke verification tests. For factory production control tests, the mandatory testing items are pressure, low velocity, constitutive law, and damping efficiency tests. In this latter testing program, the pressure test shall be performed on 100% of the devices while all the dynamic tests shall be performed on 1 every 20 units. Table 18 of EN 15129 (CEN, 2009) lists all these testing items, while further details of them are given in Section 7.4 of this standard code (CEN, 2009).

It must be highlighted that, in any case, laboratory tests must provide adequate information about the variability of the performance, depending on the operational temperature and on the frequency of excitation. It is also worth specifying that all qualification tests should be conventional and carried out according to standard protocols. Another important requirement, of the European standards (CEN, 2009), is that protection devices must be CE-marked: this mandatory conformity marking guarantees that all of the common European quality standards for materials, components, and subassemblies are compliant with code requirements (CEN, 2009). Moreover, it certifies that the production and manufacturing process is qualitycontrolled according to EN ISO 9001 (2008).

7.2.1.5 Device testing tools Several testing approaches are available to perform dynamic tests for distributed dampers. Generally speaking, tests can be performed in three conditions. One condition is that the damping device is tested without interaction with the

479

CHAPTER 7 Testing, inspection, and maintenance

building structural system (i.e., damper is tested independently from the structure). The other condition can be related to test the damping device while it is installed within the building system. The last condition is called hybrid testing in which the damper part is tested experimentally and the main structural part is numerically analyzed using a computer. In the following, the most common testing tools under the different conditions are described briefly.

7.2.1.5.1 Dampers tested independently from building structure One of the most common testing tools for damping devices is the hydraulic actuator. This is often utilized to drive sine wave motions or cyclic loads. For example, Montgomery and Christopoulos (2015) and Gong and Zhou (2016) conducted cyclic loading tests by considering various strain levels, frequencies, temperatures, and repeated loading to evaluate the performance of VE dampers. The following considerations can be drawn for different device categories: • •

•

For normal size dampers (e.g., viscous damper) and reduced-scale prototypes, commercially testing facilities are usually available at manufacturer factories. For large damping devices, giant hydraulic test benches may be proposed. The cost of this tool is high and requires costly power source. Alternatively, a drop hammer test may be employed using various drop heights to measure and verify the required damping function. For wall viscous dampers, a special testing device may be used to measure the mechanical characteristics of the damper (see Fig. 7.1 (Lu et al., 2008)). In

Steel cantilever beam

Displacement transducer 180

IA Strain gauge VWD Digital thermometer

Reinforced concrete reaction wall

5 5 580

Actuator 580

480

420

Screw

IA

A-A section

FIGURE 7.1 Test actuator for a viscous wall damper placed within frame system. With permission from Lu, X., Zhou, Y., Yan, F., 2008. Shaking Table Test and Numerical Analysis of RC Frames with Viscous Wall Dampers. J. Struct. Eng. 134 (1), 64
76.

7.2 Preinstallation Tests and Quality Control

this case, the damping device can be positioned between two beams, while it is fixed at the bottom beam it is free to move at the top beam. The actuator applies cycling loads to the top of system. Displacement transducers are installed at the top beam level to measure relative displacement of the damper.

7.2.1.5.2 Testing tools for damper attached to the building structure In addition to test applied directly to the damping devices, the performance of distributed dampers installed on the building can be investigated. In this case, the most important testing tools are briefly reviewed, as follows: •

•

Wind tunnel test. This test can be performed to investigate the performance of dampers in mitigating wind-induced responses. For this aim, the buildingscaled model equipped with reduced-scale damping devices (e.g., viscous wall dampers) can be tested in a wind tunnel. In order to measure the response, piezoelectric accelerometers can be installed in a number of specific floors (e.g., Yeung and Pan, 1998). Shaking table test. In this test, the whole damped structure mounted on the table is tested to measure both the structural system and damping system responses under simulated earthquake excitations. This type of test is alternatively called large-scale testing approach. The scaling factor of building
damper specimen depends on how large is the shaking table. To measure absolute acceleration, accelerometers should be installed at each floor level. After conducting the tests of damped building, it is appropriate to repeat the same specimen tests without dampers in order to evaluate the efficiency of dampers (Lu et al., 2008).

In literature, there can be found a lot of examples of low-rise building being tested on a shaking table. For instance, a 5-story building equipped with four common distributed dampers (oil damper, viscous damper, VE damper, and steel damper) was tested (Kasai et al., 2010) at E-Defense (NIED, Miki, Japan) that is one of the largest 3D shaking tables in the world. The most important responses of both building and dampers were measured under a real-simulated earthquake (the JR Takatori 1995 ground motion). Moreover, shaking table tests were performed by different researchers for VE dampers incorporated in building structures (e.g., Chang et al., 1992, 1995; Aiken et al., 1993; Lai et al., 1995; Gong and Zhou, 2016).

7.2.1.5.3 Testing tools for hybrid systems When rate-dependent effects are important, real-time testing is crucial. Therefore substructure (real-time) hybrid testing approach could be an appropriate strategy (Williams and Blakeborough, 2001). In this kind of testing, the damper part is tested experimentally and the main structural part is numerically analyzed using a computer program. Then, the results obtained from the two parts can be numerically assembled at each time step of the simulation. The simulated response (displacement) is often applied to the damper using servohydraulic actuator(s). For a

481

482

CHAPTER 7 Testing, inspection, and maintenance

FIGURE 7.2 A schematic illustration of hybrid test for distributed damping systems. With permission from Christenson, R., Lin, Y.Z., Emmons, A., Bass, B., 2008. Scale Experimental Verification of Semiactive Control through Real-Time Hybrid Simulation. J. Struct. Eng. 134 (4), 522
534.

better understanding, Fig. 7.2 illustrates schematically the concept of substructure hybrid test for distributed dampers. For loading the damper (e.g., oil damper) in the substructure hybrid test, in addition to hydraulic actuators and shaking tables, an inertia force driven loading (IFDL) system (Fig. 7.3) can also be employed. This system mainly consists of a concrete slab, damper support, power board and controller, rubber and roller supports, and an active mass driver as the actuator (Toyooka, 2002; Iemura et al., 2003; Iemura et al., 2006). The damper under test is positioned beneath the concrete slab attached to a rigid element; therefore the damper displacement is equal to the mass displacement (Iemura et al., 2006). The IFDL system was developed to permit an economical and accurate loading condition for energy dissipation devices (De Silva, 2007).

7.2.2 MASS, ACTIVE, SEMIACTIVE, AND HYBRID DAMPING SYSTEMS For mass damping systems, testing is crucial to verify the device performance and design parameters. For evaluating damper parameters, such as natural frequency, it is important to guarantee well tuning of the control system to the building (Cho et al., 2012). Generally speaking, no particular testing programs are specifically presented in standard codes for tuned mass dampers. Therefore the designer should advice which testing shall be conducted to establish the damping system and verify its main characteristics. According to preinstallation testing methods found in literature (discussed in Section 7.1), some useful recommendations are presented in the following, for different types of damping devices.

7.2 Preinstallation Tests and Quality Control

FIGURE 7.3 Test setup of IFDL system. Adapted from De Silva, C.W., 2007. Vibration Monitoring, Testing, and Instrumentation. Taylor & Francis. CRC Press, Boca Raton, FL.

7.2.2.1 Tuned mass dampers A series of tests could be conducted to confirm the TMD design parameters, for example, natural frequency, damping ratio, friction factor, restoring force, and other dynamic characteristics (Cho et al., 2012). These tests may include: free vibration, harmonic vibration tests (Kim et al., 2008; Cho et al., 2012; Lin et al., 2014), and even human-powered vibration tests (Cho et al., 2012). Harmonic vibration tests should cover more accurately the frequencies in the vicinity of natural frequencies of the damper (Min et al., 2014). Based on the responses from these tests the following parameters can be determined: • • •

Natural frequency of TMD using the spectrum analysis of the response Supplemental damping ratio via a logarithmic decrement method or halfpower approach (Cho et al., 2012) Noise during system movement (due to friction between steel materials in sliding-type components of dampers) and fail/safe lock mechanism (considered for more safety under a severe excitation) (Lin et al., 2014)

There could be some limitations while testing TMDs. For instance, if the moving mass of TMD is composed of concrete, the accurate testing of the device is hardly practical due to difficulties in concrete casting in factories (Cho et al., 2012). Moreover, in the case a damper prototype test is required by the designer,

483

484

CHAPTER 7 Testing, inspection, and maintenance

appropriate scale reduction factors may be used if test machine is subject to size limitations. For example, Lin et al. (2014) proposed a scaling factor 1/100 that was used for the mass of a TMD in a tall building.

7.2.2.2 Tuned liquid dampers The experimental testing of TLCDs for tall buildings can be used to not only verify the design parameters but also to identify the optimal geometry of devices (Cammelli et al., 2016). It is almost impossible to conduct a test for TLDs/ TLCDs with the use of an actual size device (Cho et al., 2012). Concerning liquid dampers, for example, TLCD, appropriate scale reductions can be used to conduct the test. For example, Cammelli et al. (2016) proposed scaling factor 1/20 that was used for testing a TLCD.

7.2.2.3 Magnetorheological dampers Experimental tests for (semiactive) magnetorheological (MR) dampers for civil engineering applications may include quasistatic tests, dynamic tests, and magnetic field tests (Yang et al., 2002, 2014). During quasistatic and dynamic tests, the main design parameters (dynamic range and damping force) of MR dampers can be calculated and then be compared with those obtained in the preliminary design (as per Section 5.4) (Xu et al., 2012; Yang et al., 2014). The quasistatic test can be carried out by using an actuator with a given servo valve (e.g., a servohydraulic controller accompanied with a hydraulic pump), and sensors to measure displacement and force (Yang et al., 2002). The facilities for the dynamic test may include an electrohydraulic servotester, a data acquisition system, and a dc power supply (Yang et al., 2014). Using these tests, the effect of changes in current, displacement, and frequency can be assessed on the force
velocity
displacement of dampers (Xu et al., 2012). The displacement level and frequency range to be used during tests can be selected in conjunction with seismic applications (Xu et al., 2012). Based on the experimental curves of force
displacement and force
velocity, the parameters of analytical models (e.g., Bouc-Wen (Section 4.1.1.2.1)), proposed for device analysis, can be calibrated as well (Xu et al., 2012). For civil engineering applications, the dynamic response time can be checked during the tests. In this regard, rapid response time is more desirable (Jeong et al., 2004; Hou et al., 2010). The magnetic field test can be adopted to evaluate the efficiency of effective damping path in MR dampers. For this aim, a device including a teslameter (a device or magnetometer to measure magnetic inductance or magnetic field strength) and a dc power supply can be utilized together (Yang et al., 2014). The effective damping path is a portion of the length between the piston and cylindrical housing in the damper that mainly contributes to the damping. Based on the magnetic field test, the results of current
magnetic induction and depth of damping path
magnetic induction can be obtained. For this aim, the increment of current (e.g., 0.2 A) and of damping depth (e.g., 2 mm) can be used (Yang et al., 2014).

7.2 Preinstallation Tests and Quality Control

7.2.2.4 Device testing tools Several testing approaches are available to perform dynamic tests for both passiveand active-isolated dampers. The most important of which are: wind tunnel, shaking table, and substructure hybrid tests.

7.2.2.4.1 Wind tunnel test Wind tunnel testing can be used to provide initial estimations of the dynamic wind load reduction provided by the damping system (Infanti et al., 2008). Moreover, wind tunnel testing can also be used to assess the structural performance of the tall building with TMDs (Xu et al., 1992; Gerges and Vickery, 2003; Infanti et al., 2008; Tuan and Shang, 2014). Some of the challenges associated with the explicit modeling of dampers within wind tunnel aeroelastic models include: space limitation within the model itself, size and weight of the actual TMD as well as stiffness and damping setup of the actual device (Kwok and Samali, 1995). Scaling laws are utilized to estimate the mass per unit length of the wind tunnel aeroelastic model, while the stiffness is typically provided by an internal metallic spine and/or by a number of metallic rods. The effect of additional damping in an aeroelastic wind tunnel model can, in some cases, be physically modeled by applying small strips of cellular or plastic tape between the different sections of the cladding of the model (Kwok and Samali, 1995; Facioni et al., 1995; Kim et al., 2008).

7.2.2.4.2 Shaking table tests Shaking table test is quite effective in evaluating the seismic performance of complex structures such as tall buildings (Jiang et al., 2014), if the scaled model structural response can represent the expected full-scale building response. The application of such test for high-rise buildings equipped with TMDs is demonstrated by some researchers, for example, Jiang et al. (2014) and Lu and Chen (2011) when the structure is responding in the linear-elastic response region. Moreover, they are also effective for TLCDs (Cammelli et al., 2016). To conduct such tests, a reduced-scale model of the building and damper is usually built where scaling ratios ranging from 1/12 to 1/50 are found appropriate. The shaking table motion can be programmed based on the solution computed from the equation of motion associated with the fundamental mode of the real building (Cammelli et al., 2016). Shaking table tests can be employed also for the performance evaluation and validation of isolated dampers (TLCDs) under wind excitations (Cammelli et al., 2016). Moreover, active tuned mass dampers (ATMDs) can be scaled and installed on shaking tables for testing. The major challenge is the proper installation of the active control system and the extrapolation of experimental results to prototypes (Kwok and Samali, 1995).

485

486

CHAPTER 7 Testing, inspection, and maintenance

7.2.2.4.3 Substructure hybrid tests In practice, substructure hybrid tests can be adopted for both passive- and active-based mass dampers, for example, TMD and ATMD, in which the damper loading experimental and numerical representations of structural response are combined (Da Silva, 2007; Heo et al., 2009). This testing technique is recommended in order to avoid constructing large size models and reducing high expenses. The substructure hybrid tests of mass dampers can be classified into following categories (Da Silva, 2007): •

•

Hybrid tests with hydraulic actuator. Hydraulic actuators are used to simulate loadings on the structure in order to create large excitation force and displacements (Tanzo et al., 1992; Igarashi et al., 1993; Igarashi, 1994; Williams and Blakeborough, 2001; Yamamoto et al., 2001). To precisely perform such tests, algorithms, for example, operator splitting numerical integration scheme (Nakashima, 1993) and compensation methods (Nakashima and Masaoka, 1999; Nakashima et al., 1999; Blakeborough et al., 2001; Horiuchi and Konno, 2001) are usually employed. Hybrid tests using shaking table. Shaking table is used to simulate loadings on experimental specimens (dampers). Tables are mostly driven themselves by hydraulic actuators; hence, actuator algorithms and technologies are directly applicable. Iemura et al. (2002) tested electromagnetic mass dampers using shaking tables. Isolated dampers, for example, TMD and TLCD, are commonly tested based on the use of shaking tables (e.g., Bairrao et al., 2008; Heo et al., 2009; Rakicevic et al., 2012; Lin et al., 2014; Min et al., 2014). This type of testing was employed for AMD and ATMD systems (Iemura et al., 1992), as well as for semiactive TMD (Lin et al., 2012) and semiactive TLCD (Yalla and Kareem, 2003).

7.2.3 BASE ISOLATION SYSTEMS 7.2.3.1 Standard code requirements In order to verify the deformation and damping characteristics of isolation systems before construction, tests are recommended by worldwide design codes. In this regard, in the United States, ASCE (2017a) suggests qualification and prototype tests. Other US standards, for example, ASCE (2010, 2013), mainly emphasize only on prototype tests. Isolation units must be examined for quality control during the tests. Their real force
displacement behavior, for adopting in the analysis of base-isolated buildings, should be determined using cycling tests of prototype isolators (ASCE, 2017a). It is also recommended to test the ability of bearings under vertical loading and cycles of shear displacement during design and maximum earthquake (ASCE, 2017a). Similar to US standards, in Europe EN 15129 (CEN, 2009) recommends two tests: type test (equivalent to prototype test) and factory production control test (equivalent to qualification test), both elastomeric isolators and sliding (pendulum) isolators. Each of these test programs may include several testing items. For

7.2 Preinstallation Tests and Quality Control

elastomeric isolators, Table 11 of CEN (2006) summarizes the relative testing items. Interested readers are encouraged to refer to Section 8.2.4 of EN 15129 (CEN, 2009) for more details. For sliding-type isolators, the type test and factory production control test are addressed in Section 8.3.4 of EN15129 (CEN, 2009) for the verification of dynamic characteristics, such as: frictional coefficient, damping capacity, and stability under repeated cycling. The Chinese standard (GB50011, 2010) recommends performing tests before installation for the verification of durability and design parameters (horizontal stiffness and effective damping ratio) of isolation systems. According to GB50011 (2010), after conducting durability tests, the stiffness and damping properties should not be altered more than 6 20% of the nominal values. When testing the design parameters of isolation systems, the vertical load limits shall be kept according to the values given by Table 12.2.3 of GB50011 (2010). For more details, readers can refer to Section 12.2 of GB50011 (2010). According to Japanese practice (JSSI Manual, 2003), before commercially fabricating all types of isolation bearings, they must be tested by a government agency in accordance with requirements similar to those recommended for prototype tests by the US codes (ASCE, 2017a; NEHRP, 2015). After this qualification, the manufacturer can present the bearings with defined characteristics, applications, and modeling instructions. In contrast, in Japanese practice, it is not mandatory to perform any project-specific test on the bearings. However, simple tests before installation of every bearing are conducted to verify its properties, similar to qualification tests recommended by the US codes (ASCE, 2017a). Moreover, the upper and lower bound of design properties are considered during testing, as stated by the US standards (Becker et al., 2015).

7.2.3.2 Base isolation testing Generally speaking, preinstallation tests of isolation systems can be integrated into four categories, as reviewed in the following sections: system characterization, prototype, qualification tests, and complete unit tests (Taylor et al., 1995; De Silva, 2007). Therein, the majority of testing program recommendations come from the US design practice, ASCE (2017a). Moreover, recommendations coming from literature are reviewed as well. Based on these testing requirements, quality control procedures are also reviewed.

7.2.3.2.1 System characterization tests These are useful to estimate the essential features of new isolation systems, for example: dependency on frequency, temperature, bilateral loading, and wear and fatigue. Moreover, the ultimate and reserve capacity for different loading conditions may be among the characteristics to be estimated. Although conducting such tests is recommended only once for a given design, they can be iterated if the design scheme is subject to significant variations (Taylor et al., 1995; De Silva, 2007).

487

488

CHAPTER 7 Testing, inspection, and maintenance

7.2.3.2.2 Prototype tests These are helpful to verify the design parameters of the isolation system before construction, such as: energy dissipation capacity, effective stiffness, effective damping, stability of hysteretic behavior, stability at peak seismic displacement, and stability under maximum and minimum vertical loads. The effect of winds on base isolation should be considered in such tests. If the isolation system behavior is depending on frequency, such tests should be conducted dynamically for a frequency range which characterizes the full-scale prototype loading rates (Taylor et al., 1995; De Silva, 2007; ASCE, 2017a). According to ASCE (2017a), fullscale prototype testing of sliding bearings is important; albeit, specimens of every bearing type with reduced scales are permitted to be tested if full-scale test is impossible. In this regard, the specimens should have the same type, material, and manufacturing quality. As recommended by ASCE (2017a), prototype tests should be performed for at least two full-size specimens (or several specimens) of each type and size of isolators. The force
displacement curve of each test specimen should be recorded for each cycling test. The cycling tests of isolation system shall be performed while designing vertical load, equivalent to the sum of the average dead load and half the influences from the live load on isolator units. A set of prototype tests can be conducted sequentially as listed below (ASCE, 2017a): 1. Twenty fully reversed loading cycles in accordance with the wind design force (for testing wind resistance). 2. The arbitrary sequence of the following items (for testing seismic resistance): a. Three fully reversed loading cycles at each of four increments of the maximum displacement, that is, 0:25DM , 0:5DM , 0:67DM , and 1:0DM (where DM is the maximum displacement of the center of rigidity of the isolation system in the direction under consideration). b. At the effective period of TM , continuously loading of one full cycle at every increments 1:0DM , 0:67DM , 0:5DM , and 0:25DM , subsequently continuously loading of one full cycle at 0:25DM , 0:5DM , 0:67DM , and 1:0DM . Between these two sequences of testing, a rest break is allowed. 3. Three fully loading cycles at the maximum displacement DM (for testing seismic resistance). 4. The arbitrary sequence of the following items (for testing seismic resistance): a. Equivalent to (30SD1 =SDS B) fully loading cycles at 0:75DM ; the number of cycles should be not less than 10. b. Above test, at the effective period of TM , including separate sets of multiple cycles; each set should include not less than five continuous cycles. If isolator units also carry vertical loads, the second test item mentioned earlier shall be performed, considering two additional vertical load cases as expressed in Eq. (5.37). Moreover, one fully reversed loading cycle shall be done at the total maximum displacement (DTM ), determined in Section 5.3.1.6. In this

7.2 Preinstallation Tests and Quality Control

regard, the combined vertical loads should be applied downward on each isolator of a common type and size, and the axial force and displacement obtained from each test should be considered as the maximum of those determined with the upper and lower bound amounts of isolation characteristics (see Chapter 5). Note that the prototype tests should be performed for different combinations of displacements along two orthogonal directions, if the isolation system behaves considerably different (e.g., higher than 15%) under unilateral and bilateral loadings (Taylor et al., 1995; De Silva, 2007; ASCE, 2017a). In this case, ASCE (2017a) recommends to conduct above-listed tests by considering the load increments along two directions as: 0:25DM and 1:0DM , 1:0DM and 0:25DM , 0:67DM and 1:0DM , and 0:67DM and 1:0DM . After conducting prototype tests, the following acceptance criteria shall be satisfied (ASCE, 2017a): • •

•

• •

•

The hysteretic plots for aforementioned tests represent a positive incremental force resisting capacity. For (3) testing item, the average postyield stiffness and energy dissipation of each cycle fall in the range between 1.05 and 0.95 of the corresponding nominal values (i.e., permissible manufacturing variation). For each displacement increment of (2) and (3) test items and for each vertical load case: • The average postyield stiffness of each cycle falls in the range between λðtest;minÞ and λðtest;maxÞ (Section 5.1.3.2.5) times the nominal value (i.e., variation in heating, rate of loading, and scragging). • The difference between effective stiffness of the two test specimens of a common type and size of the isolation and the average effective stiffness is not higher than 15% for each cycle of test. In the (4) testing item, for each specimen, the initial effective stiffness and damping over test cycles do not change more than 20%. In the (4) testing item, the average postyield stiffness and energy dissipation of each cycle fall correspondingly in the range between λðtest;minÞ and λðtest;maxÞ (Section 5.1.3.3.3) times the nominal values. The stability of all specimens of vertical load-carrying elements of the isolation system is satisfied.

7.2.3.2.3 Qualification tests Qualification tests are related to materials or members (e.g., elastomeric bearing and sliding systems), which construct the isolation unit. Concerning elastomeric bearings, tests should be carried out on hardness, bond strength, compression set, tensile strength, and elongation at break, low-temperature properties, high-temperature aging, and ozone resistance. For sliding-type members, surface roughness, characteristics of sliding interfaces, and bearing pad attachment should be tested. According to ASCE (2017a), full-scale qualification testing of sliding bearings is important.

489

490

CHAPTER 7 Testing, inspection, and maintenance

7.2.3.2.4 Complete unit tests Completed unit tests are used for testing sustained compression to verify the quality of the contact between the bearing and the steel/metal for laminated elastomeric bearings and also sliding bearings. Moreover, combined compression and shear can be tested to verify if real effective stiffness and supplemental damping are close enough to design values. ASCE (2017a) recommends to test 100% of isolators in the combined effect of compression and shear under displacements not less than (2/3)DM.

7.2.3.2.5 Quality control procedures According to ASCE (2017a), the specifications about quality control procedures (e.g., acceptable variations in tested design properties) are under responsibility of the designer. Moreover, the overall cost and schedule requirements related to conducting multiple tests for quality control should be taken into account. In this case, those testing items are crucial for isolation bearings that are straightforwardly relevant to the verification of main design characteristics. For instance, information from vertical load
deflection tests has not a great importance compared to those of lateral load
deflection tests. After conducting quality control
related tests (as described in previous sections), some inspections are needed to verify the testing results. The requirements of inspection procedures in such tests are different depending on the type of isolation system. In particular, it is recommended to refer to ASTM D 4014 (ASTM, 2012) that provides requirements for elastomeric bearings (ASCE, 2017a). Moreover, Da Silva (2007) suggested to visually review the bond of rubber and steel, surface cracks, laminate situation, and permanent deformation for laminated elastomeric bearings. In the case of sliding-type bearings, the bond of liner and metal, scoring of the stainless steel plate, leakage, and permanent deformation should be reviewed as well. Note that any bearing not satisfying the requirements should be rejected. Moreover, ASCE (2017a) recommends to check that the mean results of nominal properties of isolators measured by quality tests are fallen in the range determined by λðspec;minÞ and λðspec;maxÞ (Section 5.1.3.3.3).

7.2.3.3 Base-isolated structure testing Since the behavior of base-isolated buildings not only depends on the isolation system but also upon the superstructure response, it can be useful to employ experimental tests to verify and calibrate the analyzed models, and to provide data for the design process. There are various testing approaches for base-isolated structures such as (De Silva, 2007): • •

Quasistatic test. Albeit this type of test is simple, a main disadvantage is that the restoring forces alternate due to strain-rate effect in isolators. Shaking table test. Such a kind of test does not represent the above defect of quasistatic test, if the specimen is real-scale size. This requires a large size table, but it is quite costly in practice. Thus the specimens should be scaled

7.2 Preinstallation Tests and Quality Control

•

•

down, resulting in less accurate outcomes. As an example, Maddaloni et al. (2016) conducted a shaking table test for base-isolated structure in conjunction with MR dampers. Online hybrid test. This testing method is suitable to achieve the response of base-isolated structures if quasistatic loading equipment and powerful computers are both available. To conduct such a test, it is recommended to reduce the whole isolated structure to a single degree of freedom system with scaling down the whole model. Substructured online test. In this technique, instead of using the whole structure used for online hybrid test, a part of that (base isolation system) is considered as the experimental substructure while the remaining part is analytically modeled (Tanzo et al., 1992).

7.2.4 TESTING EXAMPLES In this part, some examples of preinstallation tests found in literature are briefly presented. The examples include both passive and semiactive devices.

7.2.4.1 Distributed damping systems 7.2.4.1.1 Viscous dampers For Taipei 101 (Taiwan, Chapter 8, Section 8.2.3), aiming at wind-based design of viscous dampers used in the TMD, cyclic loading tests were performed to evaluate the performance of dampers and confirm design parameters; moreover, a sinusoidal energy dissipation test was conducted to verify the damper capacity in sustaining temperature increase due to maximum windstorm (velocity) (e.g., Infanti et al., 2008). These tests are similar to those ones required by ASCE (2017a). Additional tests were (other than code-recommended ones) performed for Taipei 101 viscous damper, such as: •

•

Pressure valve relief test, useful to check if the relief pressure valve (for limiting damper pressure) is activated once the force in damper exceeds the peak (design) damping force. Proof pressure test was carried out to verify pressure vessel strength and seal integrity (Infanti et al., 2008). Each device must be pressured up to a certain minimum pressure (proof pressure), and be held under this condition for a given time, for example, 180 seconds (Taylor, 2011).

7.2.4.1.2 Viscoelastic dampers The experimental verification of VE coupling dampers installed within shear wall systems in tall buildings has been recently presented by Montgomery and Christopoulos (2015). The use of such damping device is in mitigating both windinduced and seismic-induced responses. Some specimens with reduced scale were tested at the University of Toronto Structural Testing Facilities in order to determine VE material characteristics. For this aim, harmonic loading tests were used

491

492

CHAPTER 7 Testing, inspection, and maintenance

with the range of strains and frequencies consistent with those of high-rise buildings. Moreover, six full-scale specimens incorporated within coupled wall systems were tested at E´cole Polytechnique in Montreal Structures Laboratory for the application in a real 50-story high-rise building in Vancouver and an 85-story high-rise building in Toronto. For the full-scale specimens, other than harmonic tests recommended by standards, tests using representative (realistic) time histories of design wind and design earthquake excitations were performed to evaluate dampers’ performance. Moreover, ultimate dynamic tests were done to validate dynamic large-amplitude features and ultimate quasistatic tests were used for the determination of ultimate strength and connection capacity of devices (Montgomery and Christopoulos, 2015).

7.2.4.1.3 Steel dampers Several full-scale tests were performed for (rotational) friction dampers at Technical University of Denmark (DTU) and at Kawakin testing facilities in Japan in order to evaluate device performances (e.g., force
displacement loops) under dynamic excitations due to earthquakes and windstorms (Mualla et al., 2012). Such dampers were examined to be installed in a 300-m high-rise building in Osaka of Japan (Abenobashi Terminal building). The tests were conducted using Instron machine (a device to enforce displacement to device) and also dynamic servotesting machine. Several parameters such as loading frequency, displacement amplitude, and number of cycles were examined during the tests (Mualla et al., 2012).

7.2.4.2 Isolated damping systems As testing examples, shaking table tests were carried out to verify the dynamic properties of TMDs for high-rise buildings (Lin et al., 2014) and of TLMD system (acting as a TMD in one direction and as a TLCD in orthogonal direction) under wind-type disturbance (Min et al., 2014). The TLMD system fabricated for a 64-story building in South Korea was tested also by Heo et al. (2009), where a real-time hybrid testing method was used, which only needs testing the damper as an experimental model.

7.2.4.3 Active, semiactive, and hybrid damping systems In literature the following examples were found for this category of dampers: •

•

A semiactive (variable) oil damper was tested by Iemura et al. (2006) using IFDL system (see Fig. 7.3); a control law was required to control the system. Cyclic (dynamic) loading tests and the determination of force
velocity
displacement were performed, as both recommended by standard codes. The performance tests of a variable (semiactive) friction damper including piezoelectric actuator (with variable voltage) were done to identify its

7.3 Commissioning and System Tuning

•

•

characteristics (Xu and Ng, 2008); moreover, semiactive controller strategies and shaking table tests were used for the device. Iemura et al. (1992) conducted a shaking table test for AMD and ATMD based on a three-DOF frame system at Kyoto University of Japan. The damper mass was considered as the mass of servomotor (to excite the system), moving mass, driving guides (for movement of TMD), and velocity meter (for measuring velocity). For MR dampers, dynamic tests were carried out using hydraulic actuators (e.g., Chooi and Oyadiji, 2008) and shaking tables (e.g., Lin and Christenson, 2011). Moreover, real-time (pseudo-dynamic) hybrid testing approach is applicable to structures equipped with such dampers (e.g., Lin and Christenson, 2011). In addition to dynamic tests, Yang et al. (2014) performed magnetic field tests in order to investigate the magnetic induction in MR dampers against the current, as well as depth of the damping path. Moreover, Yant et al. (2014) developed a testing device specifically for such tests based on sinusoidal cycling input. From these tests, the mechanical behavior, damping force
displacement hysteretic curves, and energy dissipation performance of MR dampers were evaluated (Hou et al., 2010; Xu et al., 2012; Yang et al., 2014).

7.2.5 SUMMARY In this section, testing processes and relative responses recommended for various control systems are summarized from Tables 7.2
7.6. Tables 7.2
7.4 are for viscous dampers, VE dampers, and steel dampers, respectively. In these tables, testing processes and relevant items are in accordance with those recommended in standard codes. Table 7.5 is dedicated to mass dampers (TMD, TLD, TLCD), active mass dampers (ATMD and AMD), and semiactive MR dampers. In this table, the testing processes, approaches, and items listed are collected from the literature. Moreover, Table 7.6 lists the test requirements of base isolation systems according to international standard codes.

7.3 COMMISSIONING AND SYSTEM TUNING In order to effectively tune damping devices, such that their efficiency is optimized, it is necessary to accurately design and represent the influence of additional damping system on the dynamics of structures (Main and Krenk, 2005). One of the crucial factors, to guarantee the successfulness in the implementation of control systems, is commissioning and system tuning. The European standard, EN15129 (CEN, 2009), recommends that the device manufacturer suppliers have trained personnel to execute the system control. In this regard, the manufacturer

493

494

CHAPTER 7 Testing, inspection, and maintenance

Table 7.2 Testing Processes for Viscous Dampers International Codes

Testing Process (In Bold) and Relevant Items

US standards (ASCE, 2017a, NEHRP, 2015; ASCE, 2010, 2013)

Prototype tests: Design windstorm cycling test, MCER-displacement-based fully reversed cycling test, basic characteristics tests (force
velocity
displacement), and dependency tests (on temperature, displacement, and loading frequency) Qualification tests: Similar testing items to prototype tests

Japanese standard (JSSI Manual, 2003)

Material test: Steel, physical/chemical properties of viscous fluid and filling material Performance test (seismic): Basic characteristics tests (force
velocity
displacement), dependency tests (on temperature, strain, frequency, and velocity), and special tests (durability, aging alternation, fire resistance, water resistance, heat resistance, and weathering resistance)

European standard (CEN, 2009)

Type test: Pressure test, low velocity test, constitutive law test, damping efficiency test, wind load cyclic test, seal wear test, and stroke verification test Factory production control test: Pressure test, low velocity test, constitutive law test, and damping efficiency test

Chinese standard (GB 50011, 2010)

Performance test: Delivery inspection and type approval tests on at least three devices

should provide all the relevant information (e.g., installation drawings and tolerances) for commissioning the system. In the following, some recommendations about commissioning and system tuning of distributed damping systems, mass damping systems (including active control systems), and base isolation systems are reviewed. Most of the recommendations are related to mass dampers, since tuning such systems to the building structure plays an important role in their efficiency.

7.3.1 DISTRIBUTED DAMPING SYSTEMS Commissioning requirements are not codified for distributed systems. Therefore standard practice requirements shall be specified in the design documents. Some possible verifications to be conducted could be:

7.3 Commissioning and System Tuning

Table 7.3 Testing Processes for VE Dampers

• • • •

International Codes

Testing Process (In Bold) and Relevant Items

US standards (ASCE, 2017a, NEHRP, 2015; ASCE, 2010, 2013)

Prototype tests: Design windstorm cycling test MCER-displacement-based fully reversed cycling test, basic characteristics tests (force
velocity
displacement), and dependency tests (on temperature, displacement, and loading frequency) Qualification tests: Similar testing items to prototype tests

Japanese standard (JSSI Manual, 2003)

Material test of VE material: Shear and adhesion performance, shear strain, and fire resistance. Performance test: Basic characteristics tests (force
velocity
displacement, shear strength, and shear strain), dependency tests (on temperature, strain, and frequency), and special tests (durability, fire resistance, moisture resistance, and fatigability tests)

European standard (CEN, 2009)

Type test: Pressure test, low velocity test, constitutive law test, damping efficiency test, wind load cyclic test, seal wear test, and stroke verification test Factory production control test: Pressure test, low velocity test, constitutive law test, and damping efficiency test

Chinese standard (GB 50011, 2010)

Performance test: Delivery inspection and type approval tests on at least three devices

Mounting accuracy and bolt tightening to verify safe fastening to the junctions. No gap between the marking of bolts and nuts. The main body of devices must be free of abnormalities and flaws (JSSI Manual, 2003; Ishikawa et al., 2004). Installation details must be developed in such a way that bending generated in the device assembly is minimized; moreover, out-of-plane stability shall be considered (Vail et al., 2004).

7.3.2 MASS DAMPING SYSTEMS 7.3.2.1 Tuned mass dampers TMD on-site installation and commissioning are quite important to obtain good efficiency of these systems (Tse et al., 2012; Wang et al., 2014). In practice, in

495

496

CHAPTER 7 Testing, inspection, and maintenance

Table 7.4 Testing Processes for Displacement-Dependent Dampers International Codes

Testing Process (In Bold) and Relevant Items

US standards (ASCE, 2017a, NEHRP, 2015; ASCE, 2010, 2013)

Prototype tests: Design windstorm cycling test, MCER-displacement-based fully reversed cycling test, basic characteristics tests (force
velocity
displacement), and dependency tests (on temperature, displacement, and loading frequency) Qualification tests: Similar testing items to prototype tests

Japanese standard (JSSI Manual, 2003)

Material test: Steel Performance test: Basic characteristics tests (force
velocity
displacement, elastic stiffness, postyield stiffness, and yield strength), dependency tests (on displacement velocity and repetitive cycles), and accumulated energy absorption performance test (low-cycle fatigue test and accumulated plastic deformation test)

European standard (CEN, 2009)

Material test: Steel and shape memory alloy Factory production control test of material: Steel and shape memory alloy Type test of devices: Evaluation of force versus displacement cycle, ramp test Factory production control test of devices: Evaluation of force versus displacement cycle, ramp test

Chinese standard (GB 50011, 2010)

Performance test: Delivery inspection and type approval tests on at least three devices

order to assure the system tuning between TMD and building at the time of commissioning, forced vibration tests can be performed on-site, using the TMD mass as the building exciter. In this case, a hydraulic pump may be employed to initially push the moving TMD mass (Roffel, 2012; Cho et al., 2012). Such tests were conducted for a 50-story building in Busan (Korea) to identify the dynamic features (natural frequency and damping ratio) of the TMD at top (Cho et al., 2012). It was observed that device frequency was slightly greater than the designed value due to the fact that the TMD was originally tuned using the structural features of the incomplete building system (Cho et al., 2012). In order to tackle detuning problems, the on-site frequency retuning of TMDs can be accomplished with the use of adjustable stiffness material, for example, pneumatic springs (Wiesner, 1979; Nagarajaiah and Varadarajan, 2005) or coil spring (Nagase and Hisatoku, 1992). For pendulum TMDs, a tuning frame (adjustable pendulum length) is often utilized, resulting in an adjustable tuning frequency (Irwin and Breukelman, 2001; Roffel, 2012). For installation purposes,

7.3 Commissioning and System Tuning

Table 7.5 Testing Processes for Mass and Active, Semiactive, and Hybrid Systems Damping Device

Testing Process (In Bold), Approaches (italic), and Items

TMD

Factory test (full-scale or reduced scale): Free vibration tests, forced (harmonic) vibration tests, shaking table tests, wind tunnel tests, substructure hybrid tests, basic characteristics tests (design parameters, natural frequency, damping ratio, friction factor, restoring force, and other dynamic characteristics), and special tests (noise due to friction and fail/safe lock mechanism)

TLD/TLCD

Factory test (reduced scale): Free vibration tests, forced (harmonic) vibration tests, shaking table tests, substructure hybrid tests, and basic characteristics tests (design parameters, natural frequency, damping ratio, restoring force, and other dynamic characteristics)

ATMD, AMD

Performance tests: Forced vibration tests, shaking table tests, substructure hybrid tests, basic characteristics tests (design parameters, natural frequency, damping ratio, restoring force, and other dynamic characteristics), and performance of sensors, actuators, and controllers

MR damper

Performance tests: Quasistatic tests, dynamic tests, magnetic field tests, shaking table tests, substructure hybrid tests, basic characteristics tests (force
velocity
displacement, design parameters, dynamic range and damping force), and dependency tests (current, displacement, and frequency)

Table 7.6 Testing Processes for Base Isolation Systems International Codes

Testing Process (In Bold) and Relevant Items

US standards (ASCE, 2017a, NEHRP, 2015; ASCE, 2010, 2013)

Qualification control processes (by manufacturer): Effects of heating due to cyclic dynamic motion, effects of loading rate and scragging, effects of variability and uncertainty in bearing characteristics, and effects of temperature, aging, environmental exposure, and contamination Prototype tests: Design wind-force cycling test, maximum displacement
based fully reversed cycling test, total maximum displacement
based fully reversed cycling test including vertical loads, basic characteristics tests (force
velocity
displacement, effective stiffness, effective damping, and other design parameters), and dependency tests (on temperature, displacement, and loading frequency) Qualification tests: Combined compression and shear test on isolator units (Continued)

497

498

CHAPTER 7 Testing, inspection, and maintenance

Table 7.6 Testing Processes for Base Isolation Systems Continued International Codes

Testing Process (In Bold) and Relevant Items

Japanese standard (MOLTI, 2000)

Government agency tests: Similar to US prototype tests Simple performance tests (similar to US qualification tests): Basic characteristics tests (force
displacement and other design parameters)

European standard (CEN, 2009)

Type test for elastomeric isolators: Capacity in compression under zero lateral displacement, compression stiffness test, horizontal stiffness and damping under cyclic deformation or horizontal, stiffness under a one-sided ramp loading, lateral capacity under maximum and minimum vertical loads, dependence of horizontal effective stiffness and damping on temperature, loading frequency, repeated cycling, and ageing, and creep test under vertical load Factory production control test for elastomeric isolators: Compression stiffness test, horizontal stiffness and damping under cyclic deformation or horizontal, stiffness under a one-sided ramp loading Type test for sliding isolators: Load-bearing capacity, frictional resistance force under service conditions, static coefficient of friction, and sliding isolation tests (dynamic coefficient of friction, damping capacity, and stability under repeated cycling) Factory production control test for sliding isolators: Property verification test (vertical load
bearing capacity and frictional resistance force under service conditions) and material testing

Chinese standard (GB 50011, 2010)

Performance tests: Durability tests and verification of design parameters (horizontal stiffness and effective damping ratio)

enough space should be considered given the large mass of standard TMDs. Moreover, the device movement shall be smooth according to building sways (Nagase and Hisatoku, 1992). Another important factor to take into consideration for on-site final tuning and commissioning is the relative cost compared to the relative damper cost. In Chapter 3.6.2 some case studies were considered while reviewing life-cycle cost analyses. The review of those examples showed that the cost associated with installation and commissioning was estimated to be in the range of 10%
20% (of the damper cost) (Wang et al., 2014).

7.3 Commissioning and System Tuning

7.3.2.2 TLCDs Similar to TMD, TLCD’s commissioning is not codified. In literature, recommendations can be found based on some examples (48-story residential tower in Vancouver (Irwin and Breukelman, 2001), Random House Tower in Manhattan (Tamboli, 2005), and a 64-story building in Incheon of Korea (Cho et al., 2012)), as follows: •

•

• • •

Vibration tests on buildings, in order to check the current natural frequencies at the time of device commissioning. The water oscillation can be used for the excitation, where the water movement in TLCD can be done with pressurization and depressurization of the tank from one side. Use the run-down test and random decrement technique based on measured responses of vibration tests in order to identify the dynamic properties, natural frequency, and damping ratio. Properly adjust the sluice gate to achieve the desired damping of the device. Properly adjust the water level in the columns to achieve the tuning frequency of interest. Seal off the top of a vertical column of TLCD, pressurize it, and then release the seal to tune the device to the oscillation. Then, based on the decay of oscillation, measure the supplemental damping of TLCD.

7.3.2.3 Active, semiactive, and hybrid systems The reliability of active control systems in buildings, as well as their efficiency in controlling response, is an important aspect for their commissioning. To achieve a great capability from an actively controlling system, its parameters should be tuned to the up-to-date dynamic characteristics of the building. In doing this, the response observation and identification of both the building and the control systems could be conducted on-site. For example, forced and free vibration tests can be used as explained in the following (Ikeda, 2009): •

Forced vibration tests. They can be conducted just before the construction is completed. Vibrators are usually installed close to the top floor to dynamically excite the building (the mass damper itself is sometimes employed as vibrator). Then, by switching the damping system on and off, the controlled and uncontrolled responses can be measured using sensors and then they can be compared together. The resonance frequencies can be searched by changing the harmonically exciting frequency of the vibrators, and the corresponding damping ratios can be evaluated by the identification of the resonance curve. The main shortcoming of forced vibration tests is the high efforts and expenses (Ikeda, 2009). An example of such tests is the Shanghai World Financial Center (Shanghai, China, Chapter 8, Section 8.2.14), a 492-m-high building controlled by an ATMD system mounted on the 92nd floor, where the damper was employed to excite the building. Since the ATMD was tuned

499

500

CHAPTER 7 Testing, inspection, and maintenance

•

to the fundamental mode, only this mode could be excited for the test; thus the characteristics of this mode were identified (Lu et al., 2014). Free vibration tests. The resonance mode frequencies and corresponding damping ratios, associated with the active control system, can be evaluated similar to the methods mentioned for forced vibration tests (Ikeda, 2009). An example of performing free vibration tests is in a 36-story building (before completion) in Tokyo equipped with hybrid mass damper. For creating free vibration conditions, first the damper excited the building and then it was let to freely vibrate (after the acceleration at roof level reached a certain level). Afterward, the damper was immediately switched on to the control phase (Nagashima et al., 2001). Test results can be analyzed through RDT. In this technique, by superimposing a recorded output response at a given floor, a quasifree vibration is obtained, resulting in the identification of natural frequencies and damping ratios of uncontrolled and controlled buildings (Ikeda, 2009).

Another important aspect of commissioning of this damper category is the successful implementation of the control algorithm. This is straightforwardly associated with the operational issues such as proper transfer and storage of data at the time of system activation, quality of data transferred, and algorithm robustness. Generally speaking, the integration between software and hardware (e.g., sensors, controllers, and force generators) must be satisfied to verify the reliability of active control system and its system tuning (Soong and Constantiou, 1994). Similar to TMDs, in Chapter 3.6.2, some case studies were reviewed that estimate the cost associated with installation and commissioning. For semiactive TMDs this could be in the range of 5%
10% (of the damper cost) (Tse et al., 2012).

7.3.3 BASE ISOLATION SYSTEMS For commissioning the isolation system, bearings should be installed so that the inspection and replacement are possible. To uniformly distribute the lateral loads between the bearings, the diaphragm below or above isolators (with tie beams among them) should be properly constructed. The isolation system should be executed so that the continuity of services, stairways, claddings, and elevators is met. The connections between the substructure and superstructure should be checked accurately to properly satisfy the transferring seismic forces (Mayes and Naeim, 2001). The free movement capacity of the bearings’ upper surface must be satisfied for the anticipated maximum horizontal displacement. In this case, the designer should inspect and verify the separation areas and corresponding components (Mayes and Naeim, 2001; BNBC, 2015). Note, the installer should record the bearings installation procedure and relative checks (EN, 2003; Marioni, 2006).

7.4 Fatigue of Dampers

7.4 FATIGUE OF DAMPERS Proper damper design (including their attachments) requires that all input scenarios over the expected life of the damper should be taken into consideration regardless of the minimal requirements that are necessary to meet applicable codes and standards. One of the most important aspects is the fatigue spectrum that the damper will be subjected to over its intended life. Each mechanical component of the damper and its attachments to the structure must be designed to be structurally adequate to maintain proper operation at the maximum specified load as well as to fatigue and serviceability inputs. Otherwise, catastrophic mechanical failure due to fatigue would result in a total loss of damping function. Seismic activity induces a low number of cycles. As a consequence, fatigue under seismic input is rarely a concern. This is in part due to the fact that required safety factors for dampers at maximum force levels are usually enough to cover fatigue inputs, especially since most of the cyclic input due to the seismic activity will be at force levels well below the maximum force specified. However, this needs to be verified for each application. In the case of incorporating dampers to tall buildings, oftentimes the governing input for fatigue considerations is wind. The number of cycles that dampers and their attachments will experience due to wind will be substantial. Ideally, each load transmitting component of the damper would be stressed below the fatigue endurance limit for its respective materials. The fatigue endurance limit is defined as that stress, below which the component will not fail regardless of the number of cycles it will experience. This ensures that fatigue will not be a problem and that costly and time-consuming fatigue analyses can be avoided. The considerations on fatigue design of energy dissipation devices and isolation systems are briefly discussed as follows, for different national standards: •

•

•

•

US design codes (ASCE, 2017a; ASCE, 2013; NEHRP, 2015). As recommended in these standards, the influence of fatigue should be taken into account in the lower bound and upper bound properties of control and isolation systems. In addition, those dampers that are subject to failure due to low-cycle fatigue should behave linearly under wind forces (ASCE, 2017a). European standard, EN15129 (CEN, 2009), which essentially addresses antiseismic devices, just provides a general statement for displacementdependent dampers mentioned afterward (Section 7.4.3). Japanese standard code (JSSI Manual, 2003) emphasizes on the fatigue assessment of distributed-type dampers under wind cycling loads. This could be also associated with base isolation systems in buildings, since some types of such dampers (e.g., viscous dampers) can be employed in conjunction with base isolation bearings. Chinese code, GB50011 (2010), addresses low-cycle fatigue in distributedtype dampers, as well as the condition in which such a fatigue does not appear in dampers (Sections 7.4.1 and 7.4.3).

501

502

CHAPTER 7 Testing, inspection, and maintenance

An additional suitable reference document, for fatigue data of various materials, is the Metallic Materials Properties Development and Standardization (MMPDS (Rice et al., 2003)). For all load transmitting components of a damper, the stress shall be held under the fatigue life. This is usually verified through the use of stress
number of cycles (S-N) diagrams for the applicable material in cases where it is not possible to be below the fatigue endurance limit. In other words, the S-N diagrams provide a reference for defining the fatigue life of a material at a particular stress level with all stress concentration effects taken into account. Generally speaking, for every type of damping systems, a complete damper specification shall define the fatigue limits, especially for those devices that are expected to sustain substantial cyclic input during their intended life. Subsequently, damper manufacturers should address fatigue life by design and shall be required to demonstrate adequate margin to avoid any failures due to fatigue. In the following, some details about fatigue in distributed-type dampers (viscous, VE, and displacement-dependent dampers) are presented addressing both code and literature recommendations. Also, a few details concerning mass dampers and base isolation systems are given, since fatigue issue is not particularly stated in standard codes, as well as the limited available discussion in the literature.

7.4.1 VISCOUS DAMPERS The fatigue life of viscous dampers and their components such as seals, connections, pistons, and bearing should be checked for durability in their device life under wind force (Smith and Willford, 2007; ASCE, 2013; NEHRP, 2015; ASCE, 2017a). Increasing the velocity exponent in viscous dampers may improve their fatigue life under regular winds. Note that in this case the efficiency of devices under low-speed wind excitations becomes limited. In any case, it is suggested to inspect such dampers to identify the signs of metal fatigue (Smith and Willford, 2007). Recommendations are also provided by Japanese standard, JSSI Manual (2003), which recommends checking the sealing of viscous dampers under wind excitations. Moreover, JSSI Manual (2003) states that it is not necessary to consider the fatigue of viscous fluid and of the filling material in viscous dampers, because their properties barely vary even under constant stresses (Tanaka et al., 2004). Requirements are also provided by Chinese standard (GB50011, 2001), which states how the decrease in the main performance of damping systems should not be higher than 10% after repeating 60 cycles at the maximum displacement. In this condition, dampers should not clearly exhibit low-cycle fatigue. According to the more recent version of Chinese code (GB50011, 2010), it is stated that the main performance degradation in velocity-dependent dampers should not be higher than 15% after repeating 30 cycles at the design velocity, at the fundamental frequency of structure.

7.4 Fatigue of Dampers

Other than proving a fatigue-proof design, good design practice requires that attention should be given to other aspects that could influence the fatigue resistance of the viscous damper design. These include surface condition, minimizing stress risers such as notches and sharp edges, and requiring processes such as shot peening, stress relief heat treatment, and adequate corrosion-inhibiting treatments. In general for tall buildings, wind tunnels can estimate the long-term fatigue loading due to wind vibrations and then one of two methods can be used to assess the fatigue behavior: (1) assess if the loading is below the endurance limit and if not (2) conduct a fatigue assessment using the loading and the S-N curves from the dampers.

7.4.2 VISCOELASTIC DAMPERS There are two types of VE material typically used for structural damping applications: 3M VE material which has a constant VE response over all displacements and a high-damping rubber material which has an amplitude-dependent elastoplastic material. In general, in addition to different material responses, they have different fatigue characteristics. Fatigue reduces the stiffness of the hysteretic response (Gong and Zhou, 2016). Fatigue is caused by cyclic deformations originated from cracks at VE material at the molecular level that gradually propagate due to subsequent cycling deformations/loads, finally leading to breakdown, that is, fatigue failure, in the material. Similar to other dampers an endurance limit can be defined and if loading is below this level fatigue is not an issue. If the material is adhered to the steel plates with a bonding agent, initial cracks can start at the VE material to steel
damper material interface or in the material itself and if the material is self-adhering, like 3M ISD-111 or ISD-111H material, the cracking cannot occur at the steel
damper material interface and occurs in the material itself. This type of failure should be considered when dampers are under cyclic deformations generated from large number of cycles, such as wind loading. In the case of VE material, generally fatigue would reduce the stiffness of the VE material, but would not affect the damping characteristics. Typically fatigue loading is not an issue for VE materials (Ishikawa et al., 2004). For example, 3M ISD-111 and ISD-111H materials are particularly inert, and require about 2,000,000 cycles, cycled to a strain amplitude of 60%, before the stiffness drops by more than 10%. The strain amplitudes in typical 3M applications are not expected to change the VE damper properties at all over the life of the building. In Japan, there have been applications dating back to 1992 where the dampers have been removed and retested after undergoing significant earthquake and typhoons and the properties have remained unchanged. In general for tall buildings, wind tunnels can estimate the long-term fatigue loading due to wind vibrations and then one of two methods can be used to assess the fatigue behavior: (1) assess if the loading is below the endurance limit and if not (2) conduct a fatigue assessment using the loading and the S-N curves from the dampers.

503

504

CHAPTER 7 Testing, inspection, and maintenance

7.4.3 DISPLACEMENT-DEPENDENT DAMPERS Displacement-dependent dampers have several conditions that represent the occurrence of fatigue fracture, such as rupture, deterioration of the peak yield strength, decrease in the energy absorption, and hysteretic loops being instable. Few fatigue resistance recommendations are available in codes and literature for this device category, such as: •

• •

•

•

•

•

EN 15129 (CEN, 2009) states that this device category is permitted to be installed in a structure if the fatigue resistance of their main components is larger than (with one order of magnitude) the number of cycles exerted in the tests. Japanese standard (JSSI Manual, 2003) states the low-cycle fatigue characteristics of steel dampers shall be tested under repetitive displacements. Chinese standard (GB50011, 2010) states that the main performance degradation in displacement-dependent dampers should not be higher than 15% after repeating 30 cycles at the design drift (displacement). In this condition, dampers should not clearly exhibit low-cycle fatigue. Deng et al. (2015) describe that the fatigue performance in shear panel dampers may be decreased due to the low-cycle fatigue damage that frequently appears close to the welded stiffener. Usami et al. (2011) performed low-cycle fatigue tests on buckling-restrained bracings (BRBs) and experimentally demonstrated that toe-finished welding method can efficiently enhance the fatigue life of such dampers with relatively small strains. Zhang et al. (2012) demonstrated that the use of low-yield-strength steel (LYS 100) for constructing seismic metallic dampers and shear panel dampers improves the low-cycle fatigue performance. Han et al. (2003) demonstrated that shape memory alloy (with nickel
titanium alloy) dampers have great fatigue resistance.

7.4.4 MASS DAMPING SYSTEMS Mass damping systems, for example, TMDs, should be designed for a long life in accordance with the design life of the building it is being considered for. The design of such systems should be performed so that they will be in continuous motion during the design life. This is because they are often operating not only at large excitations, but also under small ones. In some cases, the damper may be subject to more than 300 million cycles. Therefore fatigue resistance must be adequately examined for each component of the device for all loading amplitudes that are possible. Moreover, stress levels in the components have to be analyzed for fatigue resistance and stress levels must be held under the endurance limit to guarantee an infinite life (Klembczyk and Breukelman, 2000).

7.5 Building Health Monitoring

7.4.5 BASE ISOLATION SYSTEMS For base isolation systems, components and connections must be designed so that no fatigue occurs during the design life. For this aim, fatigue tests (e.g., highcycle tests and rotating-bending fatigue test) can be conducted on the isolation components. The cycle tests can be those recommended by standard codes (see test items by ASCE (2017a)). Some of the major recommendations for standard isolation devices are describe as follows: •

•

Lead dampers and lead rubber bearings show a sufficient fatigue resistance under cyclic loading due to the influence of recrystallization at ambient temperature (Pan et al., 2005; De Silva, 2007; Morita et al., 2014). In general, the fatigue resistance in natural rubber bearing is found higher than in synthetic elastomers (De Silva, 2007). U-shape (steel) damper isolators fatigue properties (e.g., the loading speed, total energy dissipation, and the number of cycles up to the fracture) can be measured through dynamic cyclic loading tests (Jiao et al., 2015). One possible fatigue indicator is the shear angle defined as the ratio between the total (horizontal) deformation and the damper height (peak-to-peak horizontal shear angle) (Jiao et al., 2015).

To further improve the fatigue resistance in isolation systems, the component surface may be coated by grease. This can be also effective to prevent fatigue crack propagation in isolator components (Morita et al., 2014).

7.5 BUILDING HEALTH MONITORING The idea of installing sensors to detect possible damage of building due to earthquake could probably track back to as early as when the blue mosque in Istanbul was built. Nowadays, the aims of deploying sensors in buildings are not only to detect the possible damage, but also to monitor its structural performance under the operational and extreme loadings including varying temperature, strong wind, and seismic loads. From an academic standpoint, the purpose of installing a structural health-monitoring (SHM) system can be classified into four different categories: 1. To monitor the environment loads—this includes the measurement of wind, weather, ambient temperature, seismic and the corrosion status of concrete, and/or the steel structures. 2. To monitor the operational loads—this includes the measurement of live and dead load acts upon the structural members being measured. 3. To monitor the structural features—this includes the measurement of static and dynamic features of the structural components.

505

506

CHAPTER 7 Testing, inspection, and maintenance

4. To monitor the structural responses—this includes the measurement of stress and strain distributions, as well as fatigue and articulation responses. Continuous monitoring of the building represents a very valuable tool for the validation of the design assumptions and to guarantee consistent performance levels of the dampers over their whole service life. In this regard, the efficiency of the control systems (especially those operating in relation with building properties) can be enhanced if the updated dynamic characteristics of the building are used to tune the damping system constantly. However, it is important to know that the dynamic properties of controlled buildings identified by full-scale measurements can be different under various excitation levels, since the operation of control system is associated with the input level. As most of the damper systems used for building structures aim to reduce wind-induced and/or seismic-induced vibrations, the deployment of sensors is mostly focused on the detection of movements as well as the extent of damage, if any, of the building structures being monitored (Xia et al., 2011, 2014, Su et al., 2013). Furthermore, it is noteworthy to mention that structural monitoring could also be very useful to verify the status of structural safety, especially during construction stage of tall buildings (Su et al., 2014). SHM is considered an integral part in most national hazard reduction programs (United States, Japan, China, etc.) (Garevski, 2013). However, code requirements are sometimes not sufficient to get enough information (Garevski, 2013). A short summary of national code and guideline recommendations is given as follows: •

•

• • •

United States: in UBC (1997), for all buildings above 10 stories (or above 6 stories with a floor area greater than 5574 m2), a minimum of three triaxial accelerographs shall be placed. The same requirements can be found in the current IBC (2018) (Appendix L) in which instruments shall be located at lowest, mid-, and top-portion of the structure. However, these instrumentation requirements are not sufficient to obtain meaningful data (Garevski, 2013). Canada: ISIS (2001) provides guidelines for SHM as a diagnostic tool. A detailed discussion about the various components of a typical SHM system (such as sensors, data acquisition, and analysis algorithms) is provided for both experts and nonexperts in the field. In addition, detailed examples are provided. Europe: SAMCO (2006) developed a guideline on existing procedures and technologies for SHM, providing recommendations on their utilization. Russia: GOSTR 53778 (2010) introduces inspection, testing technologies, and condition-based classification schemes for different structural types. China: GB-50982 (2014) describes SHM requirements for super high-rise structures as well as large span, bridge, isolation, and crossing construction structures. This code describes most of the field monitoring methods and sets the corresponding technical parameters (Yang et al., 2017).

Given the importance of tall building health monitoring, some general requirements for passively damped (distributed and mass), actively damped, and

7.5 Building Health Monitoring

base-isolated buildings are explained in the following. Moreover, case studies of high-rise building being health-monitored are presented at the end of this section.

7.5.1 DISTRIBUTED DAMPING SYSTEMS Unlike shaking table tests (see Section 7.2.1.5), that use many sensors for recording responses, a limited number of sensors (accelerometers) are usually employed in real measurements of building global response. As a rule of thumb, at least the top floor and the closest floor to the ground level shall be instrumented to record top and base accelerations, respectively (Kasai et al., 2012a,b). With the use of this, the floor acceleration amplification factor can be identified. Since the sensors frequently record accelerations, the structural displacement can be numerically calculated using the recorded acceleration responses. To this end, two methods are recommended by Kasai et al. (2012a,b). The first approach is to conduct the double integration of acceleration with high-pass filtering in the frequency domain. The second approach is to first identify the modal properties (natural frequency, damping ratio, and participation factor) of the building using the transfer function of the acceleration measured from the top to the base. Then, by adopting a modal history analysis (often for the first three modes), with the recorded base acceleration as input, the displacement time histories can be obtained.

7.5.2 MASS DAMPING SYSTEMS In tall buildings controlled by mass dampers, for example, TMDs, health-monitoring strategies and requirements could be almost identical to those explained earlier for buildings with distributed dampers. The sensors to measure building responses may be attached to multiple floors, with the floors supporting the damper preferably equipped with sensors. In addition, free and forced vibration tests can be conducted (the damper itself may be used to excite the structure) after building completion to identify the dynamic characteristics of the buildings/dampers and to detect damage on the basis of measured data analyzed (Fu and Johnson, 2014). These tests can be done as a part of health-monitoring program of tall buildings and are in addition to those conducted for commissioning and system tuning at the time of damper installation (Section 7.3.2).

7.5.3 ACTIVE, SEMIACTIVE, AND HYBRID DAMPING SYSTEMS When dealing with active, semiactive, and hybrid dampers, it is important to understand how to relate the monitoring system to the controlling system. Indeed, both systems can use the same sensor system, data acquisition, and transmission system. The identified responses of both the building system and controlling system should be evaluated in order to guarantee the effectiveness of dampers (Ikeda, 2009).

507

508

CHAPTER 7 Testing, inspection, and maintenance

The cooperation between active, semiactive, and hybrid dampers and monitoring system may be illustrated as follows (Ikeda, 2009): •

•

Observed response under earthquake excitation. In order to monitor the control system efficiency during an earthquake and guarantee the postevent damper/building life, it is possible to compare the controlled frequency transmissibility and its uncontrolled counterpart under earthquake accelerations. Note that after recording the controlled response due to an earthquake, the control system should be held switched on until the next earthquake. Then, it can be switched off to record the uncontrolled response during that earthquake. This is a disadvantage of this approach. Observed response under wind excitation. To monitor the damper performance during windstorms, the controlled response measured under wind excitation can be compared with the corresponding uncontrolled response observed under a different wind load. This is acceptable since dynamic characteristics (e.g., velocity, direction, and power) of two different wind excitations are almost independent of whether the structure is controlled or uncontrolled.

7.5.4 BASE ISOLATION SYSTEMS Monitoring systems adopted for real high-rise base-isolated buildings could be placed on both the superstructure and the substructure. The first category of sensors is often used to measure the structural accelerations on the selected floors; while the second can be applied to measure the interstory drift (displacement) between the isolation floor and the superstructure, as well as the displacement of isolators and other components (Sato et al., 2008; Kasai et al., 2012a,b). Additionally, sensors can be employed to record the uplift of the isolation system, as it becomes an important issue for taller isolated buildings.

7.5.5 CASE STUDY EXAMPLES The majority of monitoring systems of dynamic modification devices in tall buildings have been reported from Japan. The Japan Society of Seismic Isolation (JSSI) established committees for the assessment of buildings with passive dampers and isolation systems, after the Tohoku earthquake occurred on March 11, 2011. An example is the US
Japan study team, organized by the Earthquake Engineering Research Institute (EERI) and the Architectural Institute of Japan (AIJ), that investigated the utilization of response control systems (Taylor, 2012). Moreover, relevant data of monitored building equipped with damping system were presented by Kasai et al. (2012a,b, 2013). In the following, some healthmonitoring case studies of tall buildings with dampers are presented.

7.5 Building Health Monitoring

7.5.5.1 Distributed damped systems A 54-story building, retrofitted with 288 oil dampers (installed along two directions), was monitored under the 2011 Tohoku earthquake (Japan) using sensors (accelerometers) installed at the 1st, 28th, and roof floors. Fig. 7.4 illustrates the distribution of the dampers and the location of sensors for measuring acceleration. The natural frequencies and damping ratios of the building were identified. Based on the identified dynamic (modal) properties, modal analysis was performed to compare the recorded responses and those obtained by the analysis. Thanks to the monitoring system, it was shown that the damage and fatigue to structural and nonstructural components were evidently decreased with the help of the dampers; moreover, the comfort of building occupants was satisfied due to the added damping (Kasai et al., 2012a,b, 2013). Other monitored buildings in Japan (under the 2011 Tohoku earthquake) were a 21- and a 41-story structure, which include steel dampers in addition to viscoustype devices (Kasai et al., 2012a,b, 2013). It was observed that the acceleration responses at the top of the buildings were still high since dampers were not fully operating under the exciting level of earthquake. More details about these monitored buildings can be found in Kasai et al. (2012a,b, 2013). (A)

54-story building: steel frame + 288 oil dampers 42 m 63 m

(B) x, y, z (RF)

216,000

x, y, z (28F)

223 m x, y, z (1F) z y

S Y

N

X

FIGURE 7.4 A 54-story building with oil dampers (A) and position of sensors (B) (Kasai et al., 2013).

509

510

CHAPTER 7 Testing, inspection, and maintenance

7.5.5.2 Mass damped systems The full-scale measurements of four high-rise buildings in Japan equipped with TLDs were illustrated by Tamura et al. (1995). Through these investigations, the equivalent damping ratio was estimated using the root mean square value of the measured damped-free oscillation at a given period. A beat phenomenon was observed in damped-free oscillations (i.e., an interference pattern between vibrations of structure and damper with slightly different frequencies) due to the back and forth energy transfer between sloshing water in TLDs and structure. To remove this problem, the use of floating particles was recommended by Tamura et al. (1995). The full-scale measurements also demonstrated the significant nonlinear dynamic behavior of the buildings with TLD; this may be an important problem when dealing with design of passive or active dampers.

7.5.5.3 Active, semiactive, and hybrid damped systems 7.5.5.3.1 Active mass damper A 24-story (114 m) high-rise building controlled by an AMD system (installed on the penthouse), constructed in Tokyo (Japan), was equipped with a monitoring system (Yamamoto and Sone, 2014). Monitoring of the building and the AMD system began in November 2010, around 7 months after the construction was completed. The AMD has a regenerative system to save energy and supply the power to the actuators. In addition to the AMD, viscous wall dampers were installed along all the stories of the building. The AMD system was designed to satisfy only human comfort under strong wind input with an 1-year return period. Instead, viscous dampers were used to improve the building resistance under strong seismic loads. Forced vibration measurements using the AMD were conducted in order to identify natural frequencies and damping ratios, while the monitoring items included the measurement of: • • •

The building accelerations along two orthogonal directions at the position of the AMD The stroke displacements of the AMD along two orthogonal directions The electric energy supplied to the AMD along two orthogonal directions

Under the 2011 Tohoku earthquake excitation, almost no records of the AMD operation were observed. However, on September 15, 2011 another earthquake stroke the building and some records were observed for both the building and AMD. Moreover, during a typhoon (10-year return period wind) in 2011, the AMD was activated and monitored for 7 hours, and the above-listed items were measured.

7.5.5.3.2 Active-tuned mass damper Dynamic features and wind-induced responses of the Shanghai World Financial Center, a 492-m-high building controlled by an ATMD system mounted on the 92nd floor, were assessed during wind events. The damping system was designed

7.5 Building Health Monitoring

and installed to work actively under wind but to operate as a passive TMD under seismic loading. Several monitoring control analyses were conducted by various researchers using forced vibration tests and ambient measurements (e.g., Shi et al., 2012; Lu et al., 2014; He and Li, 2014). According to He and Li (2014), the building monitoring was done by considering the efficiency of the control system under typhoons. The site for the measurements was chosen about 36.8 m below the position of the ATMD. Therein, two low-frequency accelerometers in orthogonal directions were positioned for the detection of structural vibrations. Fig. 7.5 illustrates the layout of the floors including the ATMDs and accelerometers. Before proceeding, the accelerometers were calibrated with an ambient test. The acceleration responses were measured during 20 hours, covering the period in which the Typhoon Muifa occurred. It was observed that the ATMD system was activated several times during testing (when the typhoon was much closer to the building site and the related wind excitation was more significant) and the resulting damping ratio (at fundamental mode) was higher at these times. According to Lu et al. (2014), a different system of sensors was used for the building monitoring. The distributed accelerometer sensors were installed at various floors (e.g., at floors 10, 50, 60, 70, 80, and 90). The modal properties (frequency and damping ratio) of the fundamental mode and those of higher modes were identified based on ambient vibration tests. Based on the identification conducted by Shi et al. (2012), the natural frequencies and damping ratios were found, respectively, smaller and larger than those identified by He and Li (2014) for this building. The reason was related to the fact that, when the monitoring

FIGURE 7.5 Plan view of the floor with ATMDs (A) and site test floor with accelerometers (B). Adapted from He, Y.C., Li, Q., 2014. Dynamic responses of a 492-m-high tall building with active tuned mass damping system during a typhoon. Struct. Control Health Monit. 21 (5), 705
720.

511

512

CHAPTER 7 Testing, inspection, and maintenance

was done under lower wind intensities, compared to a typhoon, the ATMD operation was much more reduced.

7.5.5.3.3 Hybrid mass damper A 36-story building equipped with HMD, built in Tokyo (Japan), was equipped with a monitoring system (Nagashima et al., 2001). The HMD consisted of a gear-type pendulum and linear actuators. Indeed, two coupled HMDs were installed at the 36th floor in order to control the vibrations induced by wind. Fig. 7.6 illustrates the position of dampers and sensors in the building. The sensors were servovelocity types used for both the control system and observation of the wind-induced vibration. The ambient monitoring was conducted under the typhoon 9810 in 1998 for a period of 4 hours. In this case, the wind profile, including the averaged turbulence intensity, gust factor, and speed, was observed through the anemometer (device for measuring wind speed used also in weather station) installed at the roof level (see Fig. 7.6). Moreover, longitudinal, transversal, and torsional components of the wind load acting upon the building were obtained with the use of the measured response of the building and the acceleration of the HMD system. The influence of control on wind loads was considered in this case. In parallel to the observed (measured) responses, using the identified mathematical model of the building structure, the same responses were simulated for the case of

FIGURE 7.6 Layout of HMD system and servovelocity sensors. Adapted from Nagashima, I., Maseki, R., Asami, Y., Hirai, J., Abiru, J., 2001. Performance of hybrid mass damper system applied to a 36-storey high-rise building. Earthquake Eng. Struct. Dynam. 30 (11), 1615
1637.

7.6 Ongoing Maintenance

uncontrolled building in order to evaluate the performance of the control system (Nagashima et al., 2001).

7.5.5.4 Base isolation systems A 20-story base-isolated building, property of Tokyo Institute of Technology, located in Yokohama (Japan) was monitored for measuring accelerations, displacements, natural frequencies, and total damping ratios (Sato et al., 2008; Matsuda et al., 2012). The isolation system is composed by rubber bearings, steel dampers, and oil dampers. Fig. 7.7 illustrates the monitoring system and location of sensors in both the superstructure and isolation floor. As shown in Fig. 7.7A, accelerometers were installed on the ground floor, 1st, 2nd, 7th, 14th, and 20th floors. To measure the displacement of the isolation components (bearings and dampers), some displacement transducers were used (Fig. 7.7B). Therein, a trace recorder was employed, combined with the other measurement devices, in order to measure the large/small interstory displacement in the isolated story located between 1st and 2nd floors (Fig. 7.7C). The recorder was fixed to the bottom of the superstructure and to the top of the substructure (isolation floor). Moreover, some displacement transducers and video cameras were installed to measure the uplift in the isolator system (see Fig. 7.7B). The input
output acceleration signals were recorded under 2011 Tohoku earthquake excitation and the data were used to compute the superstructure displacement, natural frequencies, total damping ratios, and modal participation vectors (Matsuda et al., 2012).

7.6 ONGOING MAINTENANCE Tall buildings are intended to have a long service life, for example, 60
100 years. Consequently, the effective operation of dynamic modification devices must be guaranteed for the same design life. For this reason, ongoing maintenance should be carefully planned after device installation if the dynamic modification device requires it. Several factors shall be taken into account while defining a maintenance plan and some of them could include the following: • •

•

A measuring platform should be installed near the damper installation place for visual inspections at any time. The intervals between inspections should be around 2
5 years or after a large-certain scale earthquake. A record of the damper condition at that time shall be done. Check the appearance condition of the dampers by visual inspection to determine if the normal operation of the dampers is interfered or limited by the objects from outside.

513

CHAPTER 7 Testing, inspection, and maintenance

20th floor

14th floor

X

7th floor Isolator uplift (2)

θz

Z

2nd floor Isolation floor 1st floor (ground)

Y θy X

Small story drift (1)

Isolator uplift (3)

Oil damper force, deformation (2)

N 0

0 ,8

15

θx

Large story drift (2 x 2)

Small story drift (2)

(A)

0 6,20

4

W

E

(B)

S

r

750

Target (fixed to superstructure) ux

Δy 750 +

Trace record device

750

Wire type disp. sensor (fixed to isolation floor)

uy

750 + Δx

514

Wire

Wire type disp. sensor (fixed to isolation floor)

Metal plate for trace recorder

Rigid support for metal plate

(C)

FIGURE 7.7 Monitoring system for base-isolated building (A) position of sensors in 3D configuration, (B) position of sensors in isolation level, and (C) trace recorded at isolation floor (Matsuda et al., 2012).

• •

All damper components should be regularly maintained and repaired (if needed) by designated personnel. The employed personnel should have appropriate certificates and should be properly trained. Violation of operation is strictly prohibited to avoid manmade shortening of the service life of the mechanical system.

Z Y

7.6 Ongoing Maintenance

•

•

During the service life of the damper, property variations should not exceed 10% of the design values, and it should be regularly inspected to ensure the working performance (GB-50011, 2010). After experiencing flood, fire, wind, or earthquake above the design level, dampers should be checked and a function test is required to be implemented. If any problem is found on the dampers, they should be repaired/replaced timely.

In the following, code recommendation features are reviewed. Subsequently, the most important aspects of maintenance for distributed-type dampers, masstype dampers, and base isolation systems are summarized. At the end, a summary table is presented listing the most important recommendations (with relative notes) about maintenance of various control systems.

7.6.1 STANDARD RECOMMENDATIONS International standards are not very prescriptive regarding these aspects and only few of them provide some recommendations. As recommended by ASCE (2017a), the designer shall establish a maintenance schedule for each type of control device to ensure its life safety. Similarly, the Japanese code (JSSI Manual, 2003) prescribes a maintenance policy as illustrated in Fig. 7.8. As seen in the figure, ongoing maintenance for passive devices is required especially if the long-term device warranty is desired. Moreover, postearthquake assessment of dampers and base isolation system components can be considered as a part of the maintenance program (JSSI Manual, 2003; Kibayashi et al., 2004; Higashino and Okamoto, 2006). This will be further discussed in Section 7.8.

FIGURE 7.8 Maintenance policy. Adapted from Kibayashi, M., Kasai, K., Tsuji, Y., Kikuchi, M., Kimura, Y., Kobayashi, T., et al., 2004. JSSI manual for building passive control technology part-2 criteria for implementation of energy dissipation devices. 13th World Conference on Earthquake Engineering. Vancouver, B.C., Canada, August 1
6, No. 2990.

515

516

CHAPTER 7 Testing, inspection, and maintenance

7.6.2 DISTRIBUTED DAMPING SYSTEMS Distributed dampers usually require little maintenance (e.g., viscous dampers (Taylor and Constantinou, 1996; Kareem et al., 1999; Taylor, 1999)). Therefore in most of the cases good engineering practice is sufficient. Some of the devices’ standard procedures are the following: •

•

•

Viscous dampers should be designed and fabricated in order to be maintenance free during their anticipated lifetime, whereas subject to expected service condition (CEN, 2009). General maintenance should in any case consist the inspections in order to check the device conditions. The major possible problem checks to be performed are the following: leaks, cracks, metal fatigue, and corrosion. If any of the mentioned problems are detected, remedies should be considered in accordance with the supplier manual (Smith and Willford, 2007; Smith, 2016). VE damping material. 3M-ISD111 and 3M-ISD111H dampers do not require maintenance. VE devices with high-hardness rubber may be subject to inspection in order to detect some potential defects such as peeling or cracking of the VE material (Ishikawa et al., 2004); however, generally tests have shown that these defects are superficial in nature and do not reduce the damper stiffness and damping properties. Displacement-dependent dampers often have low cost and maintenance features (Pall and Marsh, 1982; Constantinou et al., 1990; Grigorian et al., 1993; Li and Reinhorn, 1995; Mualla and Belev, 2002; Pall and Pall, 2004; Lee et al., 2008; Jia, 2015). Regarding friction dampers, Pasquin et al. (2004) state that they do not require regular inspection and maintenance. For such dampers, the resistance to corrosion and maintenance can be improved using brass material (Grigorian et al., 1993). Similarly, BRBs and shear panel dampers have low-maintenance requirements (Nakashima and Chusilp, 2003).

7.6.3 MASS DAMPING SYSTEMS Mass damping systems should operate for a long service life. Therefore periodic maintenance is essential for such devices. Maintenance of TMDs is required once any detuning (off-tuning) of device (with respect to building natural modes) is detected by periodically adopting ambient measurements or forced vibration tests (Roffel, 2012). However, during maintenance, their operation should be usually stopped for retuning for several days. In this condition, the structural response may exceed the permitted level; this could be a shortcoming of maintenance for such dampers. However, maintenance-free dampers without these concerns are available (Roffel et al., 2011; Raffol, 2012; Klembczyk and Breukelman, 2000). The maintenance schedule of TMDs, established by (two) Japanese TMDs construction companies (as discussed by Tse et al. (2012)), could be as follows:

7.6 Ongoing Maintenance

• • •

Six-month warranty after device installation; the repair and replacement of broken components is included for free. Three-month (periodic) regular inspection for 2 years after device installation; this is to assess the performance of damper. One-year (periodic) inspection for 30 years after device installation to check hardware.

It is worth mentioning that the first two items listed above are also recommended by (two) Chinese companies for the maintenance program of TMDs and viscous dampers as well (Wang et al., 2014). The maintenance requirements of TLDs/TLCDs are less compared to TMDs, due to their less complex mechanism (Tamboli, 2005; De Silva, 2007; AshasiSorkhab et al., 2017). Ongoing maintenance may include inspection of water level, for example, every 3 months (Tamboli, 2005), to assure that it is almost at the same height established by the manufacturer; for this aim, a marker can be simply attached. Moreover, the components installed inside the TLCD (e.g., louver blades) to dissipate energy should be exercised as a maintenance schedule to assure the well-functioning of the device (Tamboli, 2005). At the end of the test, their position should be reset to the original position determined by the supplier.

7.6.4 ACTIVE, SEMIACTIVE, AND HYBRID DAMPING SYSTEMS This category of dampers has much higher requirements for ongoing maintenance given the need for external power source to operate. The major recommendations for the different damper devices are the following: •

Active control systems. Maintenance of these devices is strongly recommended. The reason is that such systems require external input energy to operate. Therefore most of the time they are not in operation, that is, standby mode. Hence, the long-term operation of such damping systems should be guaranteed using long-term experimental confirmation and conducting regular maintenance during their life cycle. In this regard, it is desired to get the suitable integration between software and hardware with regard to the true understanding of their interaction (Soong and Constantiou, 1994). Sensors are one of the main operative parts of active control systems. In addition, they can be simultaneously employed for the identification and diagnostics of structural performance. Hence, their existence in actively damped buildings can be important for long-term maintenance, for example, detection of the inefficient operation of device due to the change in structural behavior (Soong and Constantiou, 1994). A system monitoring (e.g., fail-safe operation methods) can be utilized to monitor the safe operation of active control systems. In this case, if the damping system is subjected to an unsafe condition, the control operation may be modified or even stopped such that subsequent problems (e.g., failing) are avoided (Soong and Constantiou, 1994).

517

518

CHAPTER 7 Testing, inspection, and maintenance

•

•

Semiactive control systems. For electrically controlled valves or mechanisms (e.g., controllable TLDs, variable-friction dampers, and variable-viscous dampers), maintenance is often difficult. In contrary, maintenance of devices based on controllable fluids (electrorheological (ER) dampers and MR dampers) is much easier. The only moving part in the latter category of devices is the piston that minimizes the maintenance requirements (Spencer and Sain, 1997). For MR dampers, a portion of maintenance and relative costs may be dedicated to the power supply needed for the activation of electromagnetic coils and then for conveying the magnetic field to the MR fluid. Hence, maintenance of such dampers can be significantly decreased if the power supply (electrical energy) is provided from the mechanical energy due to vibration, that is, the use of a self-powered MR damper with an energyharvesting device (Choi and Wereley, 2009). Hybrid dampers. As stated by Kareem et al. (1999), the active part of hybrid dampers becomes activated once the building is subjected to high-level excitations, and it is inactivate under nonsignificant loadings (i.e., the passive portion of damper performs in this condition). For this reason, as mentioned for active dampers, the active part may be out of operation for a long period of time. Therefore maintenance of hybrid dampers seems to be necessary, especially regarding the active-based portion.

7.6.5 BASE ISOLATION SYSTEMS In base-isolated buildings, the designer should determine a periodic maintenance plan for such systems and their connections (CEN, 2009; Higashino and Okamoto, 2006; De Silva, 2007). The plan shall include the replacement of the isolation system, if the isolation life is lower than the main structure life (De Silva, 2007). Two kinds of inspections can be generally carried out for isolation bearings (EN 1337 (EN, 2003)): •

•

Regular inspection is usually carried out simultaneously with the inspection of main structure and may include check of sufficient capacity for residual movement, conditions of corrosion/dust protection, various visible deficiencies, and oil leakage of dampers (if used). Principal inspection should be at bigger time intervals compared to the regular inspection and it may include those items considered in the regular inspection but with more accurate details. It is recommended to do the first principle inspection within the first year since being in service; moreover, subsequent inspections are needed after seismic events.

Note that after each type of inspections, it should be decided what actions to be taken: nothing, further inspection, or repair/replace, depending on the results obtained from the inspection. In relation with maintenance, isolation system components should be also protected against fire, corrosion, and chemical/biological reactions (Jurukovski and

7.8 Postextreme Event Inspections

Rakicevic, 1995; De Silva, 2007). For example, in case of friction pendulum bearings, in general, low maintenance is required and they can be significantly protected against corrosion by using the coated Teflon on the stainless steel (Cheng et al., 2008).

7.7 MAINTENANCE COST Part of the total cost of the control system is related to warranty and maintenance. In this case, a limited number of cost analysis of control devices has been conducted in literature (Tse et al., 2012; Wang et al., 2014), as already revised in Section 3.6.2. Some examples of maintenance costs, reviewed in detail in Section 3.6.2, are the following: • •

•

•

Viscous dampers have a maintenance cost in the range of 5%
10% of the total cost of fabrication and installation of the dampers (Wang et al., 2014). TMDs have higher maintenance costs than TLDs/TLCDs; thus it is recommended to perform maintenance once it is necessary. As reported by Wang et al. (2014), the maintenance cost of the passive TMD is between 5% and 10% of the total cost of TMD on an average. Active systems, for example, ATMDs, require higher level of maintenance costs in comparison with passive control devices (De Silva, 2007; Jackson and Scott, 2010). Hybrid dampers, for example, HMD, have lower maintenance compared to active dampers (Kwok and Samali, 1995; Kareem et al., 1999). According to Tse et al. (2012), the maintenance cost of the semiactive TMD range approximately between 5% and 10% of the total cost of damper, that is equivalent to about 0.2% of the overall construction cost of building. In general, this maintenance costs compromise the following operations (Wang et al., 2014; Tse et al., 2012): • Six months warranty • Regular system check every 3 months for the first 2 years after installation • Annual hardware check for 30 years

The maintenance requirements, with relative aspects, for the various kinds of damping systems are summarized in Table 7.7.

7.8 POSTEXTREME EVENT INSPECTIONS Tall buildings may be subject to severe loadings, during their design life, due to earthquake or windstorm (typhoon). Therefore it is recommended to inspect building components after the occurrence of any event, especially those playing a major role in the overall performance (Taylor, 1999; Smith and Willford, 2007). Hence, it is important to perform inspection of the control systems in damped tall buildings to guarantee their acceptable operation especially for subsequent events. As a general requirement, it is appropriate to furnish the architectural details (in the design documents).

519

520

CHAPTER 7 Testing, inspection, and maintenance

Table 7.7 Summarized Maintenance Issues for Different Control Systems Control System

Requirements

Inspection

Notes

Viscous dampers

Little maintenance

Leaks, cracks, metal fatigue, and corrosion




Viscoelastic dampers

No maintenance

Cracking or peeling of VE material




Steel dampers

Little maintenance

Cracks, metal fatigue, and corrosion

Use of brass material to lessen maintenance requirement

TMD

Significant maintenance

Components and repair/replacement of broken ones Possible detuning of device, e.g., using ambient measurement/forced vibration test

Off-operation of TMDs during maintenance

TLD/TLCD

Lower maintenance (than TMDs)

Periodically inspection of water level in TLDs/TLCDs, e.g., every 3 months Inspection and resetting any supplementary components, e.g., louver blades, inside TLD/TLCDs




Active isolated dampers

Significant (mandatory) maintenance

Inspection of components and repair/replacement of broken/failed ones

Utilization of control sensors for maintenance (selfidentification), e.g., detection of the inefficient operation of control device during long period Monitoring the safe operation of control system, e.g., using fail-safe operation methods (Continued)

7.8 Postextreme Event Inspections

Table 7.7 Summarized Maintenance Issues for Different Control Systems Continued Control System

Requirements

Inspection

Notes

Semiactive/ hybrid dampers

Considerable maintenance

Inspection of components and repair/replacement of broken/failed ones

Maintenance of semiactive dampers based on controlled fluid (MR dampers) is less than those that are based on electrically controlled valves

Less maintenance of hybrid dampers than active dampers

Self-powered MR damper with an energy-harvesting device requires less maintenance Base isolation

Considerable maintenance Maintenance plan by designer for system and its connections Protection of isolation system components against fire, corrosion, and chemical/biological reactions

Regular (periodic) and principal inspection

Less maintenance for friction pendulum bearings than other bearing types Use of coated Teflon on steel of friction bearings to lessen maintenance

Codes do not usually provide recommendations for postextreme event inspection. Only the Japanese standard (JSSI Manual, 2003) clearly states the necessity of postearthquake inspection of damping systems to verify the proper operation of device and to detect possible damage in the systems. Moreover, damaged devices should be replaced in case their functionality has been deteriorated (according to the maintenance schedule diagram illustrated previously in Fig. 7.8). In the following, some information about postevent inspections of distributed dampers, mass dampers, and base isolation systems is briefly explained based on available literature. The majority of information is given for base isolation components, since some references for postevent inspection are available.

7.8.1 DISTRIBUTED DAMPING SYSTEMS The following recommendations for distributed dampers have been found in literature: •

Viscous dampers. It is important to confirm that the damper is yet full of viscous fluid. Moreover, instructions may be established sometimes to

521

522

CHAPTER 7 Testing, inspection, and maintenance

•

•

•

randomly remove a few devices from the building and conduct performance tests to evaluate them. Note that the technician who inspects, repairs, or replaces the damper should be personnel of the same damper manufacturer (Taylor, 1999). Inspections conducted after some major earthquakes, for example, 2011 Tohoku of Japan and 2003 Colima of Mexico, showed the effective performance of viscous dampers (with no sign of damage to devices) to guarantee the life safety, as well as damage control of buildings (EERI, 2012; Takewaki et al., 2013). VE dampers. The postextreme event inspection for VE coupling beam elements in tall buildings is facilitated by removing the precast concrete or steel panel installed on the top of dampers (Montgomery, 2011). Friction dampers should be inspected in order to ensure that they are almost returned to their original alignment, that is, preevent position (Verganelakis and Pall, 2004). Yielding metallic dampers are often required to be replaced after strong earthquakes because of their major low-cycle fatigue problems (Liang et al., 2012).

Examples of inspected postevent behavior of distributed dampers can be found in EERI (2012), Taylor (2012), and Kasai et al. (2012a,b).

7.8.2 MASS DAMPING SYSTEMS Postextreme event inspection of mass damping systems can be considered similar to what mentioned for distributed dampers, meaning that the systems and corresponding components should be inspected to detect any deficiencies after major earthquakes and/or windstorms. Moreover, their functionality should be verified using possible on-site methods, for example, vibration tests. It must be verified that the damper properties, for example, natural frequency, are still tuned to those of buildings after extreme events. For liquid-type dampers, the designed level of water in containers, as well as position of any components embedded to improve damping of the devices, shall be inspected. Regarding the active control systems, the powered components and control devices should be inspected after extreme loadings, as well as the power supplier should be checked (Smith and Willford, 2007). An example of postevent behavior of isolated dampers is a 27-story high-rise building in Chile, Parque Araucano (Naeim et al., 2011). The building contains two TMD systems on the top to mitigate the lateral displacements under earthquakes. The TMDs are composed of several reinforced concrete boxes, filled with metal balls, which hang from the roof using a series of steel chains. The inspection after February 27, 2010 offshore Maule, Chile earthquake, showed that the TMD system operated very well, and no damage was detected in the building and in the relative control system (Naeim et al., 2011).

7.8 Postextreme Event Inspections

7.8.3 BASE ISOLATION SYSTEMS JSSI performed a series of postearthquake surveys to evaluate damage and the structural performance of seismically base-isolated buildings. Such a survey was based on the reports provided by design/construction companies and device manufacturers. In accordance with the reports, the main failures reported were relative to lead dampers embedded into base-isolated buildings, but not to the main superstructures (Kasai et al., 2013). Moreover, inspection of joint covers attaching the isolation floor to superstructure was helpful to detect any nonstructural damage (EERI, 2012). Some examples of postevent behavior of base isolation systems are: •

•

Lead dampers. They are often used as supporting components in base isolation systems and possible damages can be detected by measuring cracks with a crack tester. This can be done as a comparative evaluation of pre- and postearthquake cracks. Fig. 7.9A illustrates the cracks caused in lead dampers after the 2011 Tohoku earthquake, Japan. Fig. 7.9B depicts the comparative crack inspection before and after an earthquake in Kanto area, showing about 4 times increase in the crack depth. In addition, such an inspection can be performed for postextreme wind loading cases. For the judgment about repair/ replacement, it is appropriate to determine the remaining capacity of dampers with postevent cracks under anticipated event excitations (Kasai et al., 2013). Steel dampers, as components of base isolation systems, postevent inspection may include the residual deformation of device bars, rotation and loosening of

FIGURE 7.9 Damage to lead dampers: (A) small cracks at postearthquake condition, (B) growth of crack depth after earthquake (Kasai et al., 2013).

523

524

CHAPTER 7 Testing, inspection, and maintenance

FIGURE 7.10 Damage to steel dampers in base isolation system (Kasai et al., 2013).

•

bolts, and flaked paint on device and components. Fig. 7.10 shows these aspects occurred after the 2011 Tohoku earthquake, Japan. It is important to check any occurrence of cracks on the yielding elements (Kasai et al., 2013). Note that these steel U dampers are easily replaced and they served their purpose during the earthquake. Rubber and sliding bearings. An 18-story building, MT office in Sendai City, supported by 26 rubber bearings and 10 flat sliding bearings, was inspected after the 2011 Tohoku earthquake, Japan. No damage was detected in the structural elements of the superstructure. A few (nonstructural) damages were just detected in some joint covers between the isolation system and the building. However, these did not have any adverse influence on the main operation of the isolations (EERI, 2012).

7.9 TRANSPORTATION, INSTALLATION, AND CARE OF WORKS This section illustrates the care of works, transportation, and construction of dynamic modification devices from a practical point of view.

7.9.1 TRANSPORTATION OF DAMPING SYSTEMS Dampers are in general sturdy and durable but they still need special care during transportation. The dampers are recommended to be placed in a wooden case for transportation. Some of the major storage and handling requirements are the following: •

The damper case should always be placed vertically to prevent the devices from damage (see Fig. 7.11).

7.9 Transportation, Installation, and Care of Works

FIGURE 7.11 Storage of wall dampers. Courtesy of Dynamic Isolation Systems Inc.

• •

•

•

•

Limit on the maximum vertical height should be considered during delivery on-site (Love, 2016). Important devices or components should be protected by skids or rubber pads or other buffet, so as to avoid the damages resulted from the collision. The spare parts should be centrally stacked and should be fixed and banded firmly by steel wire ropes. Plastic rubbers should be placed between the components and the steel wire ropes. In the process of transportation, in order to prevent the coatings from damage, soft material should be placed at the component binding or fixing place to protect the components. The devices should be coated with anticorrosive materials. Any surface damage to the anticorrosive materials may cause corrosion to the devices, which should be repaired by recoating the anticorrosive materials. The devices should be placed in a relatively dry environment.

7.9.2 STORAGE/INSTALLATION OF DAMPING SYSTEMS The procedure required to install a damping system should follow standard practice routine and should involve care of works during/after installation. Some of the possible considerations to take into account are the following: • • • •

Installation admissible error should refer to the possible tolerances in the installation of the damping systems. The components must enter the site only after passing the acceptance. The components should be stacked neatly, so as to avoid deformation and damage. The components should be placed steadily on the sleeper and classified by the serial number and installation sequence.

525

526

CHAPTER 7 Testing, inspection, and maintenance

• • • • •

The storage area of the components should have good drainage to prevent water corrosion. The components that are not used should be covered by rain-proof protection. During the lifting and installation, collision and thump should be avoided. Avoid welding-supporting facilities on the components to avoid affecting the parent metal. Deliver dampers “just in time” in order to install them simultaneously to the erection of the structural system (Love, 2016).

7.9.3 CARE OF WORKS OF DAMPING SYSTEMS According to technical parameter, structural form, overall dimension, components of the dampers, there are several corresponding measures for the care of finished and semifinished products before, during, and after construction. Care of work principles can be classified as: •

•

• •

Precontrol. The arrival plan of material and equipment should be in accordance with the construction plan. The subcontractors (including the assigned subcontractors) should provide in advance the material arrival plan. This should take into account the actual schedule and the available space. The material and equipment arrival time shall be coordinated to prevent a too long time piling. Site material and equipment protection. After the materials, semifinished products, and equipment arrive to the site, they should be correctly and carefully stored. Moreover, subcontractors should be in charge of the specific care of works. Handover. The handover procedures should be carefully planned, especially between civil works (e.g., water and electricity) with different subcontractors. Crossover working care. When the subcontractors are carrying out construction, they must report to the project management in the written form about when they need to touch other professions’ finished products. In this case, the project management should assign workers to assist the subcontractors’ construction. After construction, other workers shall restore the finished products. If any subcontractor starts construction without authorization and causes damage of other products, the subcontractor should be responsible for the compensation.

7.10 RESILIENCE-BASED EARTHQUAKE DESIGN INITIATIVE AND UNITED STATES RESILIENCY COUNCIL RATING SYSTEMS Building performance is an important aspect especially for what regards protection against hazards (e.g., earthquake and windstorms). As it was shown in Chapter 3, a

7.10 Resilience-based Earthquake Design Initiative

building can be designed to satisfy different performance objectives depending on the level of the acceptable risk. For understanding the implications of this design philosophy, two different rating systems have been developed in the last years to classify building performance. In the following, the two major rating systems available in United States, REDi (2003), and USRC (2015) are briefly reviewed.

7.10.1 RESILIENCE-BASED EARTHQUAKE DESIGN INITIATIVE RATING SYSTEM Generally speaking, code-based seismic design of buildings is mainly established based on the life safety of occupants during design earthquakes. Therefore if a performance-based design approach is utilized, a certain level of damage is expected regardless of the performance level chosen (as discussed already in Section 3.5). Direct losses (e.g., repair/rebuilding costs) due to damaged buildings after a major earthquake are measurable. However, indirect losses (e.g., loss of people life, economy, and culture) are more difficult to be estimated and they could be even worse than the direct ones. Moreover, it is not definitely clear how rehabilitation of damaged buildings is economically feasible, and how buildings can be effectively reoccupied. As a consequence, it can be argued that neither code-based design nor performance-based design approaches offer promising strategies since different damage levels are allowed in design (REDi, 2013; Almufti and Willford, 2014). Resilience-based earthquake design initiative (REDi) rating system is developed to tackle the aforementioned shortcomings of the current design methods, providing some guidelines, criteria, planning, and verification techniques (REDi, 2013). Indeed, this approach offers a tool for communicating earthquake risks to engineers and building owners, and how these risks can be mitigated using integrated multidisciplinary design and planning, leading to the fast recovery of the building after large earthquakes. In other words, the REDi rating system is a design/rating tool, which satisfies the better performance of buildings following enhanced design and planning against code-based and performance-based design approaches. Three different rating tiers are proposed in the REDi (2013) approach: platinum, gold, and silver. Each of these tiers includes resilience objectives aiming at the reduction of earthquake risks in comparison with codes (see Table 7.8 for an example (REDi, 2013; Almufti and Willford, 2014)). This table lists the three rating tiers and the corresponding objectives for downtime (reoccupancy and functional recovery), direct losses, and occupant safety.

7.10.1.1 Criteria, planning, and assessment REDi (2013) states that, for every rating tier mentioned (platinum, gold, and silver), it is required to satisfy mandatory criteria in three different resilience categories: organizational, building, and ambient resilience (see illustrated criteria of

527

528

CHAPTER 7 Testing, inspection, and maintenance

Table 7.8 REDi resilience objectives for design earthquake [REDI, 2013] Downtime Rating

Reoccupancy

Functional Recovery

Direct Financial Loss

Occupant Safety

Platinum Gold Silver

Immediate Immediate ,6 months

,72 hours ,1 month ,6 months

,2.5% ,5% ,10%

Injuries unlikely Injuries unlikely Injuries possible

FIGURE 7.12 General concepts of the REDi approach (REDi, 2013).

each category in Fig. 7.12). These are the criteria necessary for the resilience design and planning. Moreover, as an evaluation task, a loss assessment study should be conducted using the FEMA’s tool (PACT (FEMA, 2012)) in order to confirm that a sufficient number of nonmandatory requirements are considered to obtain the resilience objectives (downtime and direct financial loss) of the rating tiers. In addition to loss assessment, another factor that is also outlined in the REDi guideline is the cost associated with the resilience-based design. For example, the cost premium in order to obtain a gold rating for a 43-story tall building in San Francisco was less than 2% compared to performance-based design (Almufti and Willford, 2014).

7.10.1.2 Resilience-based earthquake design initiative and response control systems The implementation of control systems (base isolation and supplementary damping systems) is one of those techniques that assists the achievement of building

7.10 Resilience-based Earthquake Design Initiative

resilience through the minimization of damage in structural/nonstructural members. Such systems permit to maintain the structure basically in elastic regime under severe earthquakes. A capacity design is established by the REDi which corresponds to base-isolated and viscously damped buildings, aiming at the increase of confidence in the performance of such buildings. Examples of criteria for building with damping system are as follows: •

•

•

For viscously damped buildings, located near fault sites, it is recommended for all the rating categories (platinum, gold, and silver) that the dampers to be designed and tested so that they accommodate the story drifts under 84th percentile MCER demands. In doing this, the 84th percentile demands given by Table C21.2-1 of NEHRP (2009) may be used, or alternatively the codebased design damper strokes may be increased by a factor of 1.9 for longperiod systems (e.g., tall buildings) (REDi, 2013). A case study tall building constructed in San Francisco, protected by megabrace systems including viscous dampers and BRBs, was designed using the REDi approach. The control system was mainly tuned for wind mitigation, but it was effective in reducing seismic-induced responses of several modes. The latter case assists to have the elastic structural system under code-based design earthquake, which is required for the achievement of the gold rating of the REDi demanded for the tall building. The benefit of building such a resilientbased design is the postearthquake immediate occupancy and functionality (Almufti et al., 2016). For base-isolated buildings, the required criteria for platinum and gold rating systems or recommended for silver rating system are (REDi, 2013): • Special braces, special walls, or special frames should be embedded in the superstructure as the lateral systems. • In the case that intermediate systems are used, the systems should be designed so that they remain elastic under the MCER demands. • Isolators should be designed and tested, providing sufficient moat clearance, to justify displacements under 84th percentile MCER demands, instead of using the code-based median MCER demands.

7.10.2 UNITED STATES RESILIENCY COUNCIL RATING SYSTEM The USRC released a manual for implementation of building rating systems for earthquake or other hazards (USRC, 2015). With the use of this, engineers and building owners are able to perform building ratings and finally review them to ensure quality control and to evaluate technical methodologies. The main objective of the USRC rating system is to better recognize the consequence of earthquakes and other major events such as hurricanes/tornadoes, floods, and blast loadings on the performance of building systems including structural, mechanical, and electrical systems, as well as architectural components (USRC, 2015).

529

530

CHAPTER 7 Testing, inspection, and maintenance

There are three independent performance measures to define for a USRC rating system: •

Safety. This rating component relates to the physical health and safety of occupants during an event (e.g., earthquake). Repair cost (damage). This corresponds to the financial loss due to repairs necessary to return the preevent condition. Functional recovery. The time to retrieve basic functioning of the building to gain its preevent use and occupancy.

• •

In general, a USRC rating system helps to obtain reliable information on the anticipated safety, repair cost, and recovery of the buildings. Table 7.9 illustrates the fundamental characteristics and general details of USRC rating system with reference to the abovementioned components. As found in this table, five rating levels, represented by one to five stars, are considered for each rating component. In this case, a higher star denotes higher building performance. Moreover, repair cost and recovery components have an additional rating level, “Not evaluated,” for which the corresponding component does not support any rating or a rating is not needed. In addition to what shown in Table 7.9, there are two possible rating types: •

Transaction rating. This kind of rating is mainly applied for financial and real estate transactions, not used for public display or marketing materials. Such a rating is limited to three stars for every three components (safety, repair cost, and recovery) listed in Table 7.9. Quality control and confirming are required for randomly selected ratings of this type using technical review by Certified Rating Professional.

Table 7.9 USRC Rating Components (USRC, 2015) Rating Level (Stars)

Rating Components Safety

Repair Cost

Recovery



Loss of life unlikely



Loss of life possible in isolated locations Loss of life likely in the building


Minimal damage: ,5% Moderate damage: ,10% Significant damage: ,20% Substantial damage: ,40% Severe damage: .40% No rating

Hours to days



Injuries and blocking of exit paths unlikely Serious injuries unlikely





Not evaluated

Days to weeks Weeks to months Months to a year More than 1 year No rating

7.10 Resilience-based Earthquake Design Initiative

•

Verified rating. This type of rating is intended for public presentation in the building entrance and in marketing materials. A technical/elevated review is needed for every rating of this kind.

The USRC rating system allows an owner to specify the desired level of performance rather than accept by default the performance of a building designed to the minimum level prescribed by the building code. For a new building, a seismic design that results in a four or five star USRC rating may add 1%
10% to construction costs, or about as much as a typical contingency budget. Major benefits of the USRC rating system can be defined as follows: • • • • • •

Performed by certified professional. Each rating is technically reviewed. Transparent—use of national standards. Credible—designed to prevent manipulation. Easy for the lay person to understand. Suitable for public display and marketing.

USRC’s approach provides consistency, usefulness, and transparency to increase free market demand for better performing buildings. The attention to building safety and business continuity will over time improve the building stock and make our cities and communities more resilient. Currently, the USRC is offering building ratings for earthquakes. Ratings for other hazards are expected to be developed soon. More information about this classification is outside the scope of this publication and interested reader should refer to USRC (2015).

531

CHAPTER

Case studies of tall buildings with dynamic modification devices

8

CHAPTER OUTLINE 8.1 Distributed Damping Systems Case Studies........................................................534 8.1.1 Columbia Tower, Seattle, Washington, United States .......................534 8.1.2 Two Union Square, Seattle, Washington, United States ....................547 8.1.3 St Francis Shangri La Place, Manila, Philippines.............................556 8.1.4 Pangu Plaza, Beijing, China ..........................................................567 8.1.5 Beijing Yintai Center, Beijing, China...............................................577 8.1.6 San Diego Central Courthouse, San Diego, California, United States...............................................................................592 8.1.7 Wuhan Poly Cultural Plaza, Wuhan, China ......................................615 8.1.8 TianJin International Trade Center, TianJin, China ...........................626 8.1.9 454 Yonge, Toronto, ON, Canada ...................................................637 8.1.10 181 Fremont Street, San Francisco, California, United States...............................................................................658 8.1.11 Atushi Building, Xin Jiang, China...................................................668 8.1.12 Costums Residential, Auckland, New Zealand .................................678 8.1.13 Connor Tower, Manila, Philippines .................................................701 8.1.14 Allianz Tower, Milan, Italy .............................................................716 8.2 Mass Damping Systems Case Studies ................................................................733 8.2.1 Citicorp Building, New York, New York City, United States ...............734 8.2.2 John Hancock Tower, Boston, Massachusetts, United States ............745 8.2.3 Taipei 101, Taipei, Taiwan.............................................................758 8.2.4 One Rincon Hill (South Tower) San Francisco, California, United States...............................................................................774 8.2.5 Comcast Center, Philadelphia, Pennsylvania, United States .............783 8.2.6 Hyatt Park Tower, Chicago, Illinois, United States ...........................788 8.2.7 Highcliff Apartments, Hong Kong...................................................801 8.2.8 Bloomberg Tower, New York City, New York, United States ...............808 8.2.9 Raffles City, Chongqing, China.......................................................818 8.2.10 L-Tower, Toronto, Canada ..............................................................828 8.2.11 14 York Street, Toronto, Canada.....................................................840

Damping Technologies for Tall Buildings. DOI: https://doi.org/10.1016/B978-0-12-815963-7.00008-7 Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved.

533

534

CHAPTER 8 Case studies of tall buildings

8.2.12 One Bloor Street East, Toronto, Canada ..........................................847 8.2.13 1151 West Georgia, Vancouver, Canada..........................................854 8.2.14 Shanghai Tower, Shanghai, China ..................................................863 8.2.15 The Independent, Austin, Texas, United States ...............................876 8.3 Base Isolation Systems Case Studies ................................................................888 8.3.1 Nunoa Capital Building, Santiago, Chile ...........................................888 8.4 Active, Semiactive, and Hybrid Systems Case Studies ........................................904 8.4.1 Thyssenkrup Test Tower, Rottweil, Germany ......................................904

Several case studies of tall buildings with dynamic modification devices have been reviewed in this chapter. A total of 31 buildings have been analyzed looking at the different aspects involved in the design not only from a structural/ mechanical point of view but also looking at the other disciplines involved in the building design (architect, consultants, mechanical, electrical, and plumbing (MEP), elevators, etc.). The information was collected from the major actors involved in the design (i.e., structural engineers, wind and damping devices consultants). The buildings reviewed are divided into the following categories: • • • •

Distributed damping systems Mass damping systems Base isolation systems Active, semiactive, and hybrid systems

Fig. 8.1 shows the location and the dynamic modification type for each of the 31 case studies reviewed. Moreover, Table 8.1 provides a summary of the major building characteristics.

8.1 DISTRIBUTED DAMPING SYSTEMS CASE STUDIES In this section, case studies equipped with distributed damping systems are reviewed. These refer to dynamic modification systems that dissipate energy through the movement between the two damper ends.

8.1.1 COLUMBIA TOWER, SEATTLE, WASHINGTON, UNITED STATES 8.1.1.1 Project data The major building data (Fig. 8.2) are summarized as follows: • • •

Year of completion: 1986 Developer: Martin Selig Real Estate Contractor: Howard S. Wright Construction Company

FIGURE 8.1 Case studies location and dynamic modification type identification.

Table 8.1 Case Study With Dynamic Modification Summary Building Function

Height (m)

Damper Type

291

Viscoelastic (VE) dampers

225.5

VE damped outrigger

2009

Office with retail podium Office with retail podium Residential

213

16 Viscous damped outrigger

China

2010

Office

191.5

Beijing

China

2007

Hotel/apartments

249.9

100 Nonlinear viscous dampers (diagonal and chevron) 1 4 VE dampers 1 36 buckling restrained braces (BRBs) 73 Viscous dampers

San Diego

2017

Courthouse

118.57

106 Viscous dampers

Wuhan

United States China

2012

Office

211.8

62 Viscous dampers (chevron and diagonal)

Tianjin

China

2014

Office

235

Viscous dampers (toggle braces)

Toronto San Francisco Xin Jiang

Canada United States China

2017 2017

Residential Residential/office

200 244

84 VE coupling dampers 32 Viscous dampers damped megabrace




Commercial/ residential Residential/retail/ office

75

56 Viscous dampers toggle brace

187

28 Viscous dampers (toggle) 1 a provisional allowance of a tuned mass damper (TMD) at the tapered apex

174.5

Viscous coupling dampers

#

Name

City

Country

Year

8.1.1

Columbia Tower Two Union Square St. Francis Shangri La Place Pangu Plaza

Seattle

1986

Manila

United States United States Philippines

Beijing

8.1.2 8.1.3

8.1.4

8.1.5 8.1.6 8.1.7 8.1.8

8.1.9 8.1.10 8.1.11

Beijing Yintai Center San Diego Courthouse Wuhan Poly Cultural Plaza TianJin International Trade Center 454 Yonge 181 Freemont Street Atushi Building

Seattle

1989

8.1.12

Customs Residential

Auckland

New Zealand

2021

8.1.13

Connor Tower

Manila

Philippines

2020

Residential with retail stores in the podium

8.1.14 8.2.1 8.2.2 8.2.3 8.2.4

Allianz Tower Citycorp Center Tower John Hancock Tower Taipei 101

8.2.9

One Rincon Hill Comcast Center Hyatt Park Tower Highcliff Apartments Bloomberg Tower Raffles City

8.2.10

Milan New York City Boston Taipei

2015 1978

Office Offices/retail

202.2 279

8 Viscous dampers Biaxial TMD with servo hydraulic actuators

1977

Offices

241

2004

Commercial/ office/hotel

508

2008

Residential

188

Dual TMD; two sliding mass blocks only in the EW direction. Tower: pendulum TMD Pinnacle: dual TMDs Tuned liquid damper (TLD)

2008

Commercial/office

296.7

Tuned liquid column dampers (TLCDs)

1999

Condominium/ Hotel Residential

257.35 252.3

Pendulum TMD, equipped with four special hydraulic dampers TLD

246

Pendulum TMD equipped with dampers

239

234

6 friction pendulum bearing with two perpendicular viscous dampers in each tower Bidirectional tuned liquid

234

Bidirectional tuned liquid

255

Bidirectional TLD

187

TLCD

New York City Chongqing

United States United States United States United States United States China

2018

L-Tower

Toronto

Canada

2015

8.2.11

14 York Street

Toronto

Canada

2015

8.2.12

One Bloor Street East 1151 West Georgia

Toronto

Canada

2015

Vancouver

Canada

2016

8.2.5 8.2.6 8.2.7 8.2.8

8.2.13

San Francisco Philadelphia

Italy United States United States Taiwan

Chicago Hong Kong

2003 2004

Condominium/ office/retail Residential/retail/ office Condominium residential/retail/ performing arts Condominium residential/retail Condominium residential/retail Condominium residential/hotel

(Continued)

Table 8.1 Case Study With Dynamic Modification Summary Continued #

Name

City

Country

Year

Building Function

Height (m)

8.2.14

Shanghai Tower The Independent Nunoa Capital Building

Shanghai

China

2016

Hotel/office

632

Austin, Texas Santiago

United States Chile

2018

Residential

209

2016

Residential

86

Rottweil

Germany

2016

Elevator test tower/office/ observation

246

8.2.15 8.3.1

8.4.1

Thyssenkrupp Test Tower

Damper Type Pendulum TMD with an eddy current damping system and a viscous damper TLD 24 Isolators: 16 lead rubber bearings (lead rubber bearing) and 8 rubber bearings (RB). hybrid mass damper (HMD) (pendulum rope supported/linear motor as actuator)

8.1 Distributed Damping Systems Case Studies

FIGURE 8.2 Columbia Center: tower building overview.

• • • • • • • • •

Architectural designer: Chester L. Lindsey Architects Structural designer: Skilling Ward Rogers Barkshire (SWRB now Magnusson Klemencic Associates (MKA), United States) Damping supplier: 3M Company Height: 284 m Interstory height: 3.5 m Stories: 76 above grade 1 7 below grade Gross area of the tower: 139,354 m2 Building function: Office with retail podium and below grades parking Structural material: Composite: ASTM A572, A36, A500, A53 steel 26
52 MPa concrete, typical 65.6 MPa encasing concrete at megacolumns

539

540

CHAPTER 8 Case studies of tall buildings

• •

Floor plan area: Information not available Damping system: Intrinsic damping: 1% Added damping: 2.5% Damping type: 260 viscoelastic dampers placed on select braces Intrinsic 10-year acceleration response: 25 mg Damped 10-year acceleration response: 14
18 mg

8.1.1.2 Introduction/history The Columbia Seafirst Center is Seattle’s tallest structure and was the tenth tallest building in the world at the time of completion. This 76-story structure, situated between Fourth and Fifth Avenues and Columbia and Cherry Streets, covers more than 1.5 million gross square feet area. It is enclosed in three stairstepping towers with contrasting concave and convex surfaces. The tower has the appearance of three slender towers of varying heights integrated into one structure with three concave and three flat surfaces. The uppermost tower is 291 m above Fourth Avenue and 320 m above sea level.

8.1.1.3 Structural system The building is framed in structural steel with a triangular core of braced moment frames. In addition, there are three composite high-strength columns, located at the corners of the triangular core (Fig. 8.3). These columns, ductile to meet earthquake requirements, measure 2.44 m 3 3.66 m in plan and contain 65.5 MPa concrete. The composite columns carry a large part of the gravity load and act, together with the braced moment frames, to resist wind and earthquake forces. Select braces are fitted with large viscoelastic (VE) dampers in parallel with diagonal braces in order to reduce accelerations at service level wind events which may cause occupant discomfort (Figs. 8.4 and 8.5). This was the first use of large VE dampers in a major building. The braced moment frames include steel arches or K-bracing. There are about 11 giant arches stacked vertically in each frame. These arches span about 47.2 m and, in some cases, have a rise which extends up to 11 floors. The arches transfer vertical and horizontal loads from the steel members to the three main composite concrete columns. Secondary trusses are located between these main high-rise arches in order to provide torsional resistance. The sway of the building is limited to H/600. The floor framing for the plaza, arcade, and parking levels is a 11.4 cm concrete slab over 7.6 m metal deck. Above the plaza level, the floor framing is 5.1 cm concrete slab over 5.1 cm deck. All steel floor beams are composite with the concrete slabs.

8.1.1.3.1 Building fundamental periods The first three building fundamental periods were both estimated and measured, as follows: • Estimated: 5.30 seconds (first mode), 4.90 seconds (second mode), 4.46 seconds (third mode)

8.1 Distributed Damping Systems Case Studies

FIGURE 8.3 Columbia Center: typical tower floor framing.

• Measured: 5.33 seconds (first mode), 4.92 seconds (second mode), 4.13 seconds (third mode)

8.1.1.3.2 Damping strategy utilized The building damping was estimated to be 1.0% of critical. In order to reduce the wind-induced dynamic response acceleration to not more than 0.020 g for a 10-year storm (annual recurrence probability of 0.10), the damping had to be increased to 3.5%.

8.1.1.3.3 Additional damping provided by the damping system The damping system was designed to provide added damping of 2.5%, for a total damping of 3.5% in each major axis (NS and EW).

541

542

CHAPTER 8 Case studies of tall buildings

FIGURE 8.4 Columbia Center: lateral force resisting system.

8.1.1.3.4 Building cost versus damping cost Damping cost was less than 0.5% of building cost.

8.1.1.3.5 Building code The governing building code for this building is the 1979 Uniform Building Code (UBC, 1979) including City of Seattle Building Code Amendments. Since occupant comfort is not addressed in the governing building code, an acceptable acceleration target of 20 mg at a 10-year recurrence interval was set based on published literature on the subject.

8.1.1.3.6 Peer-reviewed project VE system reviewed by 3M Company. Comprehensive wind effects reviewed by A. Davenport, University of West Ontario, Canada.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.5 Columbia Center: braced moment frame system.

8.1.1.4 Damping overview 8.1.1.4.1 Damping strategy Accelerations due to excitation of a 10-year wind event were determined by wind tunnel testing and found to be higher than the acceptable level established. To reduce these accelerations, supplemental damping was required; the amount of damping needed was also established by the wind tunnel study. Two-hundred and sixty VE dampers were placed on the main lateral system braces to provide additional damping rather than increasing member sizes to stiffen the building. This resulted in a building with satisfactory accelerations using less material than would have been necessary otherwise. The dampers consist of VE material

543

544

CHAPTER 8 Case studies of tall buildings

FIGURE 8.6 Columbia Center: VE damper typical scheme.

FIGURE 8.7 Columbia Center: VE damper installed in the building.

sandwiched between steel plates, which dissipate energy as the plates move relative to one another due to wind. The braces chosen to have dampers were those expected to have sufficient deformation to effectively activate the damper for energy dissipation. Fig. 8.6 shows a diagram of the dampers and Fig. 8.7 shows photographs of the braces with dampers installed in the building, demonstrating the varying size of the dampers.

8.1 Distributed Damping Systems Case Studies

8.1.1.4.2 Damping type Two-hundred and sixty linear VE dampers were installed alongside structural braces in the structure.

8.1.1.4.3 Structural and damping design The building was linearly analyzed for strength using SWRB STRUDL, one of the earliest finite element computer software. The analysis was performed on an IBM 370 computer. Modal analysis was also performed with the model to find the fundamental periods and mode shapes. Due to computational limitations, the effects of supplemental damping were evaluated using hand calculations. The level of total damping required, 3.5%, was determined in the wind tunnel study. Although there is not a permanent monitoring system, the building was temporarily instrumented in 1986 and the fundamental periods were verified. Code and guidelines. UBC (1979) was followed. At that time, there were no specific guidelines applicable for VE damper design. Analysis modeling and software. The building was linearly analyzed using SWRB STRUDL software (finite element program) with modal analysis. The dampers were designed by hand calculations. Design principles. The dampers were designed using the theory of linear viscoelasticity, using hand calculations. Design phase considerations. Schematic design: Wind tunnel tests showed additional damping would be necessary—SWRB (now MKA) did schematic designs for possible locations of VE dampers. VE dampers were chosen from a previous firm experienced in designing VE dampers for the World Trade Centers. During DD phase the location alongside certain brace members of the braceframe lateral system that provided adequate displacements was identified. CD phase started with the collaboration with 3M Company to develop a VE material with the correct properties for the project.

8.1.1.4.4 Architectural integration strategy No architectural integration strategy was required because the dampers were located alongside already present structural brace member, hidden within walls.

8.1.1.4.5 MEP integration strategy No MEP integration was necessary based upon the damper locations alongside the structural braces. No interaction between dampers and MEP systems.

8.1.1.4.6 Elevator and other mechanical devices integration strategy Elevator and other mechanical devices did not need to be considered with the type and location of the VE dampers chosen. There was no interaction between elevators or other mechanical devices and the dampers.

545

546

CHAPTER 8 Case studies of tall buildings

8.1.1.4.7 Experimental tests Dampers were tested for QA/QC and temperature effects on MTS testing system at Portec Co., Chicago. Early testing, as well as acceptance testing of the completed dampers, was performed to verify design properties of VE material.

8.1.1.4.8 Monitoring system SWRB (now MKA) installed and monitored accelerometers for a number of years after installation of dampers. No significant storms were recorded, but some building properties and wind tunnel findings were confirmed.

8.1.1.4.9 Maintenance strategy The VE dampers require no maintenance.

8.1.1.5 Comparison of different design strategies Due to space constraints for using a tuned mass damper (TMD) type damper, only VE type dampers were investigated.

8.1.1.5.1 Structural options considered It was not economically feasible to increase the stiffness of the building enough to reduce the accelerations. The only viable economical solution was to add damping.

8.1.1.5.2 Damping solution considered Once it was determined that added damping was required by wind tunnel testing, based on building configuration and space available, VE dampers were the only type of damping solution considered. Experience from the World Trade Center project done by the firm some years back allowed a quicker decision to use VE dampers.

8.1.1.5.3 Cost
benefit analysis The VE damping system was chosen in early stage of design because of cost saving and probable reliability of achieving the desired added damping required. Damping cost. Damping cost was less than 0.5% of building cost.

8.1.1.6 Lesson learned and recommendations 8.1.1.6.1 Difficulties in the design During early design, testing, and manufacture of prototype dampers, some issues were encountered with the bonding of the VE material to the steel plates. These issues were solved and the schedule was met; SWRB (now MKA) also had to design the testing program for QA/QC from scratch.

8.1.1.6.2 Design innovative solutions First application of VE dampers alongside lateral system braces.

8.1 Distributed Damping Systems Case Studies

8.1.1.6.3 Possible improvements Improvements in the VE material itself with regard to its energy absorbing properties that reduce when the temperature increases would be an improvement. Also, having a preinstalled monitoring system could give beneficial information.

8.1.2 TWO UNION SQUARE, SEATTLE, WASHINGTON, UNITED STATES 8.1.2.1 Project data The major building data (Fig. 8.8) are summarized as follows: • • • • • • • • • • • •

• •

Year of completion: 1989 Developer: Unico Properties Contractor: Turner Construction Architectural designer: NBBJ Structural designer: Skilling Ward Rogers Barkshire (SWRB now Magnusson Klemencic Associates (MKA), United States) Damping supplier: 3M Company Height: 225.5 m Interstory height: 3.66 m Stories: 56 above grade 1 4 below grade Gross area of the tower: 139,355 m2 Building function: Office with retail podium and below grade parking Structural material: Composite: ASTM A572, A36, A500 steel 26
52 MPa concrete, typical 131 MPa in concrete filled pipe megacolumns Floor plan area: Information not available Damping system Intrinsic damping: 1% Added damping: 1.5% Damping type: 16 large high-efficiency VE dampers placed at the connections between outrigger braces and outrigger columns Intrinsic 10-year acceleration response: 28 mg Damped 10-year acceleration response: 19 mg

8.1.2.2 Introduction/history Two Union Square is a 56-story multiuse office building. This award-winning project was the most economical building of this height ever built at the time. The project emphasizes column-free space on 50 floors and unobstructed views from 10 corner offices on each floor. Incorporating the most advanced application of a composite system for its time, the building was the

547

548

CHAPTER 8 Case studies of tall buildings

FIGURE 8.8 Two Union Square: building overview.

first to utilize steel pipes filled with a world-record-breaking high-strength 131 MPa concrete. The structure also incorporates the first hyperefficient VE dampers, with only 16 required in the entire 56-story building. The innovative structural design shortened construction time and reduced structural costs by 36%.

8.1.2.3 Structural system The building’s gravity system is structural steel with concrete on metal deck floors (Fig. 8.9). The lateral system consists of a braced moment-resisting frame

8.1 Distributed Damping Systems Case Studies

FIGURE 8.9 Two Union Square: typical tower framing plan.

core which utilizes large diameter concrete filled pipe columns with high-strength concrete (see Fig. 8.10 for full lateral system diagram). At levels 35 through 38, outrigger braces connect the core frames to concrete filled pipe outrigger columns to the north and south (NS) (Fig. 8.11). At the connection of the braces to the outrigger columns, VE dampers are used to dissipate energy and increase the buildings damping (Fig. 8.10).

8.1.2.3.1 Building fundamental periods The first three building fundamental periods were estimated as follows: 6.25 seconds (first mode), 5.97 seconds (second mode), and 4.22 seconds (third mode).

8.1.2.3.2 Damping strategy utilized The basic building damping was estimated to be 1.0% of critical. In order to reduce the wind-induced dynamic response accelerations to not more than 0.020 g

549

550

CHAPTER 8 Case studies of tall buildings

FIGURE 8.10 Two Union Square: lateral system.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.11 Two Union Square: outrigger braces at L35
L38.

for a 10-year storm (annual recurrence probability of 0.10), the damping had to be increased to 2.5%.

8.1.2.3.3 Additional damping provided by the damping system The damping system was designed to provide added damping of 1.5% for a total of 2.5% in each major axis (NS and EW).

8.1.2.3.4 Building cost versus damping cost Damping cost was less than 0.25% of building cost.

8.1.2.3.5 Building code The governing building code for this building was UBC (1982) with City of Seattle Amendments. Since occupant comfort was not addressed in the governing building code, an acceptable acceleration target of 20 mg at a 10-year recurrence interval was set based on published literature on the subject.

8.1.2.3.6 Peer-reviewed project VE system reviewed by 3M company. Comprehensive wind effects reviewed by A. Davenport, University of West Ontario, Canada.

8.1.2.4 Damping overview 8.1.2.4.1 Damping strategy A wind tunnel study performed at the University of Western Ontario indicated that service level winds would cause accelerations higher than acceptable, so 16 hyperefficient VE dampers were placed between the outrigger braces and columns

551

552

CHAPTER 8 Case studies of tall buildings

FIGURE 8.12 Two Union Square: VE damper.

to provide additional damping rather than increasing member sizes to stiffen the building. This resulted in a building with satisfactory accelerations and reduced construction costs by 36%. The dampers consist of VE material sandwiched between steel plates (Fig. 8.12), which dissipate energy as the plates move relative to one another due to wind or earthquake and deform the material. In order to allow enough relative motion to activate the dampers, a built-up steel pipe inside another pipe, encased in the concrete filled tube outrigger column separated by a layer of lubricating oil (essentially creating what is now a buckling restrained brace, BRB), was used as a “spring” between the stiff brace and the stiff outrigger columns. Fig. 8.12 shows a diagram of the dampers and Fig. 8.13 shows 3D model of the damper. Figs. 8.14 and 8.15 show the damper and brace during installation, respectively.

8.1.2.4.2 Damping type Sixteen high-efficiency VE dampers.

8.1.2.4.3 Structural and damping design The building was linearly analyzed for strength using SWRB STRUDL. Modal analysis was also performed with the model to find the fundamental periods and mode shapes. Due to computational limitations, the effects of supplemental damping were evaluated using manual calculations. The level of total damping required was determined in the wind tunnel study. The benefits of the dampers were not considered when designing for strength and are only used to control accelerations due to wind.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.13 Two Union Square: 3D model of damper.

Code and guidelines. UBC (1982), no other specified guidelines were followed. Analysis modeling and software. The building was linearly analyzed using SWRB STRUDL (finite element program) with modal analysis. The dampers were designed by manual calculations. Design principles. The dampers were designed using the theory of linear viscoelasticity, using manual calculations. Design phase considerations. Schematic design: Wind tunnel tests showed additional damping would be necessary—SWRB (now MKA) did schematic designs for several possible locations of VE dampers. During DD phase the location of the outriggers columns was identified as a good location for large displacements. At CD phase a thicker high-efficiency VE material was developed to accommodate large displacements.

8.1.2.4.4 Architectural integration strategy No architectural integration strategy was required because the dampers were located within a column.

553

554

CHAPTER 8 Case studies of tall buildings

FIGURE 8.14 Two Union Square: damper mid-installation.

8.1.2.4.5 MEP integration strategy No MEP integration was necessary based upon the damper locations within the outriggers columns.

8.1.2.4.6 Elevator and other mechanical devices integration strategy Elevator and other mechanical devices did not need to be considered with the type and location of the VE dampers chosen.

8.1.2.4.7 Experimental tests Dampers were tested for QA/QC and temperature effects on MTS testing system at the Boeing Co. Early testing, as well as acceptance testing of the completed dampers, was performed to verify the design properties of the VE material.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.15 Two Union Square: brace, damper, and outrigger column, during installation.

8.1.2.4.8 Monitoring system No monitoring system was installed in the field. Owner did not opt to have the dampers monitored.

8.1.2.4.9 Maintenance strategy The VE dampers require no maintenance.

8.1.2.5 Comparison of different design strategies Based on the desire to not use space for a TMD type damper, only VE type dampers were investigated.

555

556

CHAPTER 8 Case studies of tall buildings

8.1.2.5.1 Structural options considered It was not economically feasible to increase the stiffness of the building enough to reduce the accelerations. The only viable economical solution was to add damping.

8.1.2.5.2 Damping solution considered Once it was determined that added damping was required by wind tunnel testing, based on building configuration and space available, VE dampers were the only type of damping solution considered. Experience from the Columbia Center project a few years back allowed a quicker decision to use VE dampers.

8.1.2.5.3 Cost
benefit analysis The VE damping system was chosen in early design because of cost saving and probable reliability of achieving the desired added damping required. Damping cost. Damping cost was less than 0.25% of building cost.

8.1.2.6 Lesson learned and recommendations 8.1.2.6.1 Difficulties in the design Finding a location which provided adequate displacement for placing the VE dampers was a challenge. In essence, a location of movement in the outrigger columns was manufactured that elongate and shorten as the building sways. This was accomplished by taking out a part of the composite column and placing a small steel pipe column in its place to take design loads but allow movement for the dampers. Furthermore, because the amount of movement of these locations, a thicker VE material was needed than ever previously used for this application. At first, degradation of damping properties with increase in temperature was an issue. Working with 3M company, variety of materials and varied damper design were considered to come up with a final solution that performed well for the desired amount of damping.

8.1.2.6.2 Design innovative solutions First application of large hyperefficient VE dampers.

8.1.2.6.3 Possible improvements Improvements in the VE material itself with regard to its energy absorbing properties, not reducing with increasing temperature, would be an improvement.

8.1.3 ST FRANCIS SHANGRI LA PLACE, MANILA, PHILIPPINES 8.1.3.1 Project data The major building data (Fig. 8.16) are summarized as follows: • • •

Year of completion: 2009 Developer/contractor: Kuok Group Architectural designer: Wong Tung International Ltd/Recio 1 Casas Architects

8.1 Distributed Damping Systems Case Studies

FIGURE 8.16 St Francis Shangri La Place: building overview.

• • • • • • • • • • • •

Structural designer: Arup (United States) Damping supplier: FIP Industriale Testing laboratory: FIP Industriale Height: 213 m—Twin Towers Interstory height: typ. 3.2 m Slenderness ratio: 6 Stories: 60 above grade Gross area of the tower: 197,473 m2 (total) Building function: Residential Structural material: Concrete classes: 69 MPa Dampers: 32 dampers, 16 in each tower, 2 per outrigger Floor plan area: 38 m square in plan

8.1.3.2 Introduction/history The St Francis Shangri La Place are twin-tower residential condominium in Manila, Philippines. They are currently the tallest twin towers in the Philippines.

8.1.3.3 Structural system St Francis Shangri La Place is a development of two similar residential buildings 210 m tall and approximately 38 m square in plan located in a region of typhoon winds and UBC Zone 4 seismic conditions. Each building comprises a central core, connected in the middle by diagonally reinforced coupling beams (Figs. 8.17 and 8.18). The thickness of these walls increases toward the base, to

557

558

CHAPTER 8 Case studies of tall buildings

FIGURE 8.17 St Francis Shangri La Place: tower 1 location of outrigger walls and axis system.

allow for increase in both axial and shear force. A supplementary moment frame provides both additional overturning resistance for the towers and gravity support for the concrete slab floor. For each of the two buildings, eight outrigger walls are attached to the core approximately half way up the building (Fig. 8.19). Two dampers are attached to the end of each of these outrigger walls, i.e., a total of 16 dampers per building.

8.1.3.3.1 Building fundamental periods The approximate fundamental vibration modes of the building are as follows (Fig. 8.20): Tower 1: 5.73 (X direction), 5.60 (Y direction) and 4.7 (torsion) seconds Tower 2: 8.80 (X direction), 5.82 (Y direction) and 4.9 (torsion) seconds

8.1 Distributed Damping Systems Case Studies

FIGURE 8.18 St Francis Shangri La Place: tower 2 location of outrigger walls and axis system.

8.1.3.3.2 Damping strategy utilized The damped outrigger system was used for both buildings. Details found later (Section 8.1.3.4).

8.1.3.3.3 Additional damping provided by the damping system The additional damping achieved in each direction for 100-year wind varied between 5.2% and 11.2% of critical for the two buildings and two principal directions. Note that the final level of damping used in the design uses reduced values to allow for potential inefficiencies in the system that are not captured in the analysis, such as flexibility in the damper system (Maxwell spring effect).

559

560

CHAPTER 8 Case studies of tall buildings

FIGURE 8.19 St Francis Shangri La Place: dampers outrigger location.

FIGURE 8.20 St Francis Shangri La Place: fundamental frequencies and mode shapes of tower 1.

8.1 Distributed Damping Systems Case Studies

8.1.3.3.4 Building cost versus damping cost This information is not available.

8.1.3.3.5 Building code The governing code was the National Structural Code of the Philippines (NSCP, 1999), which is very similar to UBC (1997). The design was also a very early example of performance-based design in the Philippines, and reference was made to Federal Emergency Management Agency (FEMA) 356 (ASCE, 2000) during the design process.

8.1.3.3.6 Peer-reviewed project The project was peer reviewed by an internationally renowned academic expert. Additionally, the client also employed a separate structural firm to review.

8.1.3.3.7 Design forces This information is not available.

8.1.3.3.8 Expected performance Wind acceleration for 10-year wind is 9 mg (without dampers 25 mg).

8.1.3.4 Damping overview 8.1.3.4.1 Damping strategy A tall building in an area of high seismicity and high winds needs to be carefully designed to ensure the adequate balance and stiffness and strength is achieved. Conventional practice is to stiffen a building in order to reduce the dynamic response under wind loading. However, this has the effect of increasing the seismic base shear that is attracted. By adding supplementary damping to the structure, it is possible to reduce the flexural stiffness of the building to minimize seismic base shear, and at the same time control the wind response. For this building, it could be achieved by the addition of dampers at the end of the outriggers. Thus, the damped outrigger design was realized for this building.

8.1.3.4.2 Damping type The damped outrigger is a modification of a standard outrigger building (Fig. 8.21). In this case the outrigger is only connected to the core, to the perimeter columns via dampers, and not to intermediate floor beams or slab. As a building undergoes dynamic sway motion, there is relative vertical motion between the perimeter columns and the ends of stiff outrigger elements cantilevering from the core. A damper is inserted across this structural discontinuity, dissipating energy during the cyclic motion, and resulting in an increase in the overall damping of the building.

561

562

CHAPTER 8 Case studies of tall buildings

FIGURE 8.21 St Francis Shangri La Place: outrigger damper (patent pending).

The location of the dampers is shown in Fig. 8.19. The dampers have been placed side-by-side for the ease of construction and because of restrictions on space (Figs. 8.22 and 8.23). Two dampers per wall were chosen, instead of one, for the following reasons: • • •

Cheaper to supply Easier to install Greater redundancy

Behavior of damper. The damper has a nonlinear characteristic described below: F1 5 C1 v2

At an upper threshold velocity, the behavior changes to

(8.1)

8.1 Distributed Damping Systems Case Studies

FIGURE 8.22 St Francis Shangri La Place: close-up of damper.

F 5 F1 1 C2 v0:1

(8.2)

This change in behavior was chosen to prevent overload caused by extreme seismic demands and is implemented by the use of a pressure release valve. The threshold force F1 is the maximum force seen in the dampers during wind loading. At low levels of damper velocity the force mobilized is very low. This reduces the working pressure within the damper, increases the lifetime of the moving components, and reduces susceptibility to fatigue. The damping system provides more damping when the wind loading approaches the maximum design levels. Damper strokes and forces. The peak stroke during the design wind storm was predicted to be approximately 50 mm. For the 2475-year return period earthquake this increased to 200 mm. The peak force was 2.2 MN per damper, with two dampers per outrigger. Damper optimization. The effective C value of the damper was adjusted to both minimize the inertial wind loads on the building and maintain a practical upper force demand in the damper.

8.1.3.4.3 Structural and damping design Code and guidelines. National Structural Code of the Philippines (NSCP, 1999) was used as the primary design code.

563

564

CHAPTER 8 Case studies of tall buildings

FIGURE 8.23 St Francis Shangri La Place: damper installation.

Analysis modeling and software. Oasys GSA was used as the primary analysis software to calculate static loads within the structure. For nonlinear seismic analysis, LS-DYNA was used to perform NLRHA. In order to calculate the level of damping to be assumed during the wind response calculations, MSC-Nastran was used to calculate dynamic response. The direct frequency method was used to calculate damping using the halfpower bandwidth method. Design phase considerations. Following initial design, wind tunnel testing was used to better quantify the wind loading. Alongside the wind tunnel tests, a

8.1 Distributed Damping Systems Case Studies

Table 8.2 St Francis Shangri La Place: Variation in Wind Overturning Moment With Damping Method

Assumption

Wind tunnel Wind tunnel Wind tunnel

Code wind speed, intrinsic damping (1.0%) Code wind speed, damping 5 7.5% Wind speed from climate study, damping 5 7.5%

Factored Overturning Moment (GNm) 7.4 4.5 3.7

Table 8.3 St Francis Shangri La Place: Variation in Lateral Accelerations for 10-Year Wind With Damping Method

Assumption

Wind tunnel

Climate study wind speed, intrinsic damping 5 1.0% Climate study wind speed, damping 5 7.5% Suggested limit for 10-year wind

Wind tunnel

Peak Lateral Acceleration (mg) 25.6 9.4 15

directional climate study was performed. This made a significant difference to the final assessment of wind-induced response. Considering only one load direction, the base overturning moments are given in Table 8.2. Instead, lateral accelerations were predicted for the 10-year wind, using wind tunnel tests as provided in Table 8.3.

8.1.3.4.4 Architectural integration strategy The dampers were placed at the end of the outrigger walls and within small access corridors, so they could be accessed for inspection and potential maintenance.

8.1.3.4.5 MEP integration strategy There was no interaction with MEP systems.

8.1.3.4.6 Elevator and other mechanical devices integration strategy There was no interaction with other elements directly, although the dampers were sized such that they could be fitted within the service elevator if they needed to be replaced.

565

566

CHAPTER 8 Case studies of tall buildings

8.1.3.4.7 Experimental tests A sample of the dampers was tested by the manufacturer, FIP Industriale. The objective of the testing was to assess the response of the units under: • • •

Typhoon conditions. The test simulated 3 hours of varying wind force. The maximum credible earthquake. Five cycles of the peak force and stroke. All test dampers performed within specification.

8.1.3.4.8 Monitoring system No monitoring system was installed during the construction process.

8.1.3.4.9 Maintenance strategy The design life of the dampers is 50 years. There is no planned maintenance.

8.1.3.5 Comparison of different design strategies 8.1.3.5.1 Structural options considered The structural system within the buildings is a “dual system.” Owing to other architectural constraints, the building was insufficiently stiff to control dynamic wind force adequately and methods of achieving sufficient stiffness proved to be uneconomic. Studies of dynamic modification devices showed that the addition of damping improved structural response for both wind and seismic forces. Thus, no further “undamped” options were looked at.

8.1.3.5.2 Damping solution considered In addition to the damped outrigger, a TMD and a tuned slosh damper (TSD) were considered. While feasible, these were not pursued, partly because of the high value placed on the floor space at the top of the building.

8.1.3.5.3 Cost
benefit analysis Damping cost. The damper units cost approximately US$1 million. When compared to options not including damper, there was a net cost saving of US $5
10 million.

8.1.3.6 Lesson learned and recommendations This was the first example of the damper outrigger configuration in a tall building. As the design progressed, a number of lessons were learned: • • •

Nonlinear dampers, with exponents less than 1, have the advantage to limit the damper force for seismic events. Linear dampers (alpha exponent of 1) give more constant levels of damping over a large range of amplitudes. The maximum achievable level of damping is not always practical since the damping force may be too high.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.24 Pangu Plaza: building overview.

8.1.4 PANGU PLAZA, BEIJING, CHINA 8.1.4.1 Project data The major building data (Fig. 8.24) are summarized as follows: • • • • • • •

Year of completion: 2010 Developer: Beijing Pangu Investment Co., Ltd (China) Contractor: Beijing Shougang Construction Group Co., Ltd (China) Owner: Beijing Pangu Investiment Co., Ltd (China) Architectural designer: C.Y LEE & Partners (China) Structural designer: John A. Martin & Associates, Inc. (United States) Damping supplier:

567

568

CHAPTER 8 Case studies of tall buildings

•

• • • • • •

•

Taylor Devices, Inc. (United States) Beijing Qitai Shock Control and Scientific Development Co., Ltd (China) Testing laboratory: Laboratory of Taylor Devices, Inc. (United States) Laboratory of Peking University for wind (China) Height: 191.5 m Interstory height: 4.6 m Stories: 40 above grade 1 5 below grade Gross area of the tower: 141,000 m2 Building function: Office building Structural material: Concrete classes: C65 Structural steel: Q345 Dampers: 67DP-17248-01;67DP-17249-01; 67DP-17290-01 Rebar: HRB335 Floor plan area: Rectangular plan having dimensions L 5 43 m 3 43 m, total floor area 1849 m2

8.1.4.2 Introduction/history Pangu Plaza is a 191-m, 39-story steel high-rise building located at Beijing close to 2008 Olympic main stadium. The building is the unique landmark complex in the core area of the 2500-acre-Olympic park. The exterior of the building is like a dragon, which combines the quintessence of China’s 5000-year culture. The building provides a wide view of Beijing Olympic venues Bird’s Nest and Water Cube.

8.1.4.3 Structural system The building has a steel frame structure system. The plane layout of the building is designed with three cores; 32 frame columns with 8 at each side of the outside cube, and the distance between columns is 6 m. The inside tube is composed of 20 columns, which is connected with the outside tube by steel beams, and the space between the tubes is 14.6 m. The core tube is composed of four columns, whose dimensions are larger than the other columns. In addition, cantilever trusses set up at the top of the building. The floor slabs adopt steel
concrete composite slabs.

8.1.4.3.1 Building fundamental periods The building fundamental vibration modes are shown in Fig. 8.25. The first mode (T1 5 5.99 seconds) is mainly flexural along the X direction, the second model (T2 5 5.19 seconds) is mainly flexural along the Y direction, and the third one (T3 5 4.32 seconds) is mainly torsional (see Table 8.4 for the mode participation factor for the first six mode shapes).

8.1 Distributed Damping Systems Case Studies

FIGURE 8.25 Pangu Plaza: fundamental frequencies and mode shapes.

Table 8.4 Pangu Plaza: Structural Periods and Mass Participation Factors Period

UX (%)

UY (%)

RZ (%)

5.993 5.191 4.323 1.847 1.659 1.466

74.281 0.005 0.005 13.283 0.003 0.002

0.003 70.123 3.681 0.001 12.639 1.373

0.010 3.889 73.719 0.006 0.966 9.559

8.1.4.3.2 Damping strategy utilized About 108 viscous dampers and 36 BRBs were used to protect the structure from an intensity 8 earthquake hazard and type III soil ((GB 50011, 2001), Fig. 8.26). Among the viscous dampers, 96 were chevron braces and 8 were diagonal braces installed in every story from level 4 to 39 (Fig. 8.27). Diagonal brace dampers were utilized at the locations where the space was very limited and the structure was slender. Additionally, at the top cantilever truss, 4 dampers (with a capacity of 1500 kN) were installed to reduce the vertical vibration caused by wind.

569

570

CHAPTER 8 Case studies of tall buildings

FIGURE 8.26 Pangu Plaza: typical framing plan.

FIGURE 8.27 Pangu Plaza: damper distribution.

8.1.4.3.3 Additional damping provided by the damping system In case the whole dissipative capacity of the dampers is exploited (at the design service temperature), the nominal additional damping is 2% for earthquake frequencies.

8.1 Distributed Damping Systems Case Studies

8.1.4.3.4 Building cost versus damping cost The 108 dampers cost f5,800,000 (US$920,000), including material, fabrication, shipping, customs duty, tax, braces, and installation fee. Configuration cost including installation, material, and fabrication of steel elements is f10,000 (US$1600) each ton. In order to obtain the comparison results for both seismic response and economic effectiveness, another traditional aseismic way of enlarging the size of structural columns and braces was computed as comparison. The plan 1 is to enlarge the sizes of columns and braces, and the sum to cost is f12,646,000 (US$2,000,000), including f9,328,000 (US$1,500,000) of column and f3,318,000 (US$530,000) of brace. The plan 2 is to enlarge size and number of the original braces, and the cost is f8,529,000 (US$1,350,000).

8.1.4.3.5 Building code Several building codes were utilized and some of them were considered as reference documents. Among all, the most important are: • • •

Code for seismic design of buildings: GB50011 (2001) Technical specification for steel structure of tall buildings: JGJ99 (1998) Guidelines for the evaluation of the response of occupants of fixed structures, especially buildings and offshore structures, to low-frequency horizontal motion (0.063
1.0 Hz): ISO 6897 (1984)

8.1.4.3.6 Peer-reviewed project None.

8.1.4.3.7 Design forces Time history analysis is adopted and the peak value of time history waves are determined as follows (Fig. 8.28 and Table 8.4): • • •

The maximum acceleration under frequency earthquake is 70 cm/s2. The maximum acceleration under design earthquake (DE) is 200 cm/s2. The maximum acceleration under rare earthquake is 400 cm/s2. 400

400

300

300

200

200

100

100

0

0

–100

–100

–200

–200

–300

–300

–400

–400

Natural record AYL1-D-X FIGURE 8.28 Pangu Plaza: time history wave diagrams.

Artificial record SYS4-D-X

571

572

CHAPTER 8 Case studies of tall buildings

FIGURE 8.29 Pangu Plaza: 1500 kN damper installed.

8.1.4.4 Damping overview 8.1.4.4.1 Damping strategy Steel structures have a very low intrinsic damping ratio especially for high-rise buildings. Dampers attached to this building can improve total structural damping and reduce the response under the wind and earthquake loads. The main goal for this building was to improve the structural performance under earthquakes. Viscous damper is a velocity-dependent device. For the same condition, the greater the damper displacement and velocity, the greater damping the damper can supply. The dampers are arranged to building locations with higher displacement demands (Figs. 8.29 and 8.30). The building model was analyzed under earthquake and wind loads, with both fluid viscous dampers (FVDs) and BRBs (or UBB), as seismic protection system. A repeated iteration procedure of design and analysis was carried for the optimization. The complete seismic response on the horizontal and vertical directions shows that the FVDs are highly effective to reduce the structural as well as the secondary system response (Table 8.5).

8.1 Distributed Damping Systems Case Studies

FIGURE 8.30 Pangu Plaza: chevron brace damper and traditional diagonal brace.

Table 8.5 Pangu Plaza: Time History Waves Name

Direction

Type

AY1-D,Z,S-X,Y,Z AY2-D,Z,S-X,Y,Z AY3-D,Z,S-X,Y,Z SYS4-D,Z,S-X,Y,Z

Mostly subordination Mostly subordination Mostly subordination Mostly subordination

and and and and

vertical vertical vertical vertical

Natural Natural Natural Artificial

Table 8.6 Pangu Plaza: Damper Parameters 1 2 3

Type

Stiff. (kN/m)

Vel. Exp. α

C (kN(s/m)α)

F (kN)

D (mm)

Qua.

FVED FVD FVD

700 No No

0.3 0.3 0.5

3000 3000 4000

1000 1000 1500

6 100 6 100 6 150

8 96 4

8.1.4.4.2 Damping type The dampers utilized have the following main properties (Table 8.6): • • • • • •

Type of damper: nonlinear viscous damper Design life: 50 years (with maintenance) Maximum damper force and stoke: F 5 1200 kN and D 5 6 100 mm Nonlinear constitutive law: F 5 Cvα, with α 5 0.4 Damping constant: C 5 2000 kN(s/m)α and C 5 1500 kN(s/m)α Main vibration period: 3 seconds , T ,6 seconds

573

574

CHAPTER 8 Case studies of tall buildings

8.1.4.4.3 Structural and damping design Code and guidelines. The major codes utilized in the design process were the following: • • •

Code for seismic design of buildings: GB50011 (2001) Technical specification for steel structure of tall buildings: JGJ99 (1998) Guidelines for the evaluation of the response of occupants of fixed structures, especially buildings and offshore structures, to low-frequency horizontal motion (0.063
1.0 Hz): ISO 6897 (1984)

Analysis modeling and software. The software ETABS (CSI, 2016a) was used to analyze the building with viscous dampers through nonlinear time history method (Fig. 8.31). The damping devices were modeled by Maxwell model (Chapter 4) in which the spring and dashpot parts are connected in series. Design principles. The design principle is to reduce the whole structure seismic response, including the story drifts (Fig. 8.32), the top displacement and acceleration, the base shear force and moment, the torsion of the structure, the response

FIGURE 8.31 Pangu Plaza: 3D structural model.

8.1 Distributed Damping Systems Case Studies

Original

Plan 1

Plan 2

Damper plan

41 36 31 26 21 16 11 6 1 0

0.0011 0.0022 0.0033 0.0044 0.0055 0.0066 0.0077

FIGURE 8.32 Pangu Plaza: story drift at X direction under moderate earthquake.

of the secondary system under frequent, design, and rare earthquake, respectively, required in Chinese seismic code (GB50011, 2001). Design phase considerations. According to the original building model, the chosen numbers of dampers were determined to satisfy the code requirements. Numbers and locations of dampers wear adjusted until the most optimal solution was achieved.

8.1.4.4.4 Architectural integration strategy Dampers were connected by diagonal and chevron braces. The damper locations were determined based on the principle of not affecting the space and architecture view. Besides, the dampers were covered by decoration with a maintenance hole.

8.1.4.4.5 MEP integration strategy There is no interaction between damping system and MEP elements.

8.1.4.4.6 Elevator and other mechanical devices integration strategy There is no interaction between damping system and elevator or other mechanical devices.

8.1.4.4.7 Experimental tests Before all dampers were installed, a number of tests were carried out. The production tests were executed at Taylor Inc. factory. The owner selected a third party laboratory and organized the designer, the constructor, and the supervisor to witness the damper performance test.

575

576

CHAPTER 8 Case studies of tall buildings

8.1.4.4.8 Monitoring system Since damper quality and efficiency can be guaranteed, monitoring system was not designed and installed.

8.1.4.4.9 Maintenance strategy Taylor Devices provided a warranty for 35 years and a maintenance-free design for the building life. All surfaces of the damper are coated with corrosion protection. No periodic maintenance, inspection, or spare parts are required for FVDs, except for rare earthquake, artificial damage, or long time work.

8.1.4.5 Comparison of different design strategies 8.1.4.5.1 Structural options considered The steel frames were designed with diagonal bracing core gravity and lateral system with outrigger truss were connected to the inner core and outside moment frames at 16th and 30th floors.

8.1.4.5.2 Damping solution considered Different damper locations and numbers were considered. According to the seismic effects and economic performance we utilized the solutions of 108 dampers.

8.1.4.5.3 Cost
benefit analysis Damping cost. The cost of 108 sets of Taylor Devices dampers including all the shipping, customs duty, Chinese tax, braces, and installation fee was f5,800,000 (US$920,000). The cost of the configuration including the material, fabrication, and installation of steel elements is f10,000 (US$1600) per ton. Design implications. In order to obtain cooperated results for both seismic response and economic effectively, other traditional seismic schemes, of enlarging the size of structural columns and braces, were computed for comparison.

8.1.4.6 Lesson learned and recommendations 8.1.4.6.1 Difficulties in the design During the design both wind and seismic forces were considered. In addition, at the top cantilever truss, vertical vibration caused by wind forces cannot be neglected during the design phase. Furthermore, given the technical difficulties related with a new type of device in Chinese construction field, it will need a long time for dynamic modification devices to be applied in practice by more and more engineers.

8.1.4.6.2 Design innovative solutions Some cross brace configurations have to be used for the limited space. Eight fluid VE dampers were used at locations where there was no space for brace but additional stiffness was needed.

8.1 Distributed Damping Systems Case Studies

Table 8.7 Pangu Plaza: Original and New Structure Dynamic Properties Plan

Original

Plan 1

Plan 2

Dampers Plan

First mode Second mode Third mode Fourth mode Fifth mode

5.99 5.19 4.32 1.84 1.65

5.82 4.99 4.19 1.77 1.57

5.05 4.39 3.37 1.50 1.36

5.99 5.19 4.32 1.84 1.65

Table 8.8 Pangu Plaza: Economic Comparison With Different Plans Item Tradition plan

Plan 1

Enlarge the sizes of columns and brace

2

Enlarge original brace size and number

Damper plan

Through install viscous dampers in the structure add the attached damping ratio for the structure, improve the capacity of resist seismic.

Direct Costs (million US dollar) Column (1.5) Brace (0.53) Sum to 2.0 Sum to 1.4 sum to 0.9

8.1.4.6.3 Possible improvements One possible improvement would be the utilization of dampers in the strengthening layers for a better damping effect (Tables 8.7
8.9).

8.1.5 BEIJING YINTAI CENTER, BEIJING, CHINA 8.1.5.1 Project data The major building data (Fig. 8.33) are summarized as follows: • • • • • •

•

•

Year of completion: 2007 Developer: Beijing Yintai Property Co., Ltd Contractor: Beijing Urban Construction Group Co., Ltd Owner: Beijing Yintai Property Co., Ltd Architectural designer: John Portman & Associates Inc. Structural designer: John A. Martin & Associates Inc. China Electronics Engineering Design Institute (CEEDI) Damping supplier: Taylor Devices, Inc. (North Tonawanda, NY, United States) Beijing Qitai Shock Control and Scientific Development Co. Ltd Testing laboratory:

577

Table 8.9 Pangu Plaza: Summary of Damper Vibration Absorption Effect Item

Analysis Case

Story drift corner

Frequency earthquake direction X Frequency earthquake direction Y Design earthquake direction X Design earthquake direction Y Rare earthquake direction X Rare earthquake direction Y Frequency earthquake direction X Frequency earthquake direction Y Design earthquake direction X Design earthquake direction Y Rare earthquake direction X Rare earthquake direction Y Design earthquake

Vertex displacement (m)

Torsion displacement ratio

Without Damper

With Damper

Effect (%)

1/850
1/1600

15
40

1/650
1/1600

15
40

1/300
1/500

12
30

1/650
1/ 1100 1/400
1/ 1100 1/220
1/ 400 1/200
1/ 450 1/80
1/200

1/300
1/550

12
30

1/120
1/240

8
15

1/60
1/180

1/80
1/230

9
16

0.146

0.122

16.4

0.117

0.099

15.5

0.496

0.444

10.6

0.268

0.235

12.1

0.670

0.593

11.5

0.705

0.636

9.8

Story 40 Story 1
15 all larger than 1.2

All story less than 1.2

20

Remark China seismic code (GB50011, 2001) requires that the maximum is 1/500

Generally less than 1/300 for super-high steel structure

China high-rise building code (JGJ99, 1998) requires that the maximum is 1/70

An important parameter

China seismic code (GB50011, 2001) requirements: if the ratio (the most flexible floor displacement/the average of the two ends displacement) larger than1.2, the structure is irregular.

Floor acceleration spectrum maximum (m/s2) Vertex vertical acceleration/ ground vertical acceleration Base shear (kN)

Base bending moment (kNm)

Base torsional moment (kNm)

35th story 40th story

6 10

3 6

50 40

An important parameter

Rare earthquake

4

2

50

An important parameter

Rare earthquake direction X Rare earthquake direction Y Rare earthquake direction X Rare earthquake direction Y Rare earthquake direction X Rare earthquake direction Y

91527

76203.38

16.7

An important parameter to measure the ability of resist seismic for the structure

151107

108130.5

28.4

3950811

3686721

6.7

3885778

3611556

7.1

3523130

2681187

23.9

4898810

3650227

25.5

Note: Frequency earthquake means the maximum acceleration is 70 cm/s2. Design earthquake means the maximum acceleration is 200 cm/s2. Rare earthquake means the maximum acceleration is 400 cm/s2.

580

CHAPTER 8 Case studies of tall buildings

FIGURE 8.33 Beijing Yintai Center: building overview.

• • • • • •

Taylor Devices, Inc. (United States) The Structural and Earthquake-Resistant Testing Center of the School of Civil Engineering at Harbin Institute of Technology (HIT), China Height: 249.9 m Interstory height: 3.3 m Stories: 62 above grade and 4 below grade Gross area of the tower: Rectangular plan having dimensions L 5 40 m 3 40 m, total floor area 113,000 m2 Building function: Boutique hotel and serviced apartments Structural material: Concrete classes: Cores and columns from story 1 to story 4: C70

8.1 Distributed Damping Systems Case Studies

•

Structural steel: Q345 Dampers: FVDs per Taylor Devices drawings 67DP-17893-01-1 and 67DP-17893-01-2 Floor plan area: Rectangular plan having dimensions L 5 24 m 3 61 m, total floor area 1464 m2

8.1.5.2 Introduction/history As the tallest skyscraper along Chang’an Avenue, Beijing Yintai Centre sits on the southwestern corner of the “Golden Cross” intersection of the China World Trade Centre Bridge in Beijing’s Central Business District (CBD). After completion, with a height of 249.9 m, Beijing Yintai Centre became a new prestigious landmark of Beijing. Its central tower in the middle reaches 66 stories and 249.9 m in height, comprising of Park Hyatt Beijing, Park Hyatt Penthouses, and Park Hyatt Residences. The central tower is flanked on each side by two symmetrical 52-story, 186-m tall superior office towers. Standing like a tripod, the three square towers resemble the Chinese character “龤” (pin), which translates to “quality”. Beijing Yintai Centre has become an exceptional commercial, recreational, entertainment, and fashion destination in Beijing.

8.1.5.3 Structural system The tower is a frame-core wall structure with levels 3 and 4 as rigid girder converting stories. Furthermore, levels 23, 33, 47, and 56 are rigid stories set up with steel trusses between the inside tube and outside tube. The four stories are composed of 12 outside frame columns and center concrete core tubes. The inside and outside tube of the top-out steel structure is a frame tube structure system. The outer tube columns are typically spaced 5 m with some being 2.5 m. Because of the vertical transportation, the layout of the inside frame tube column spaces are unsymmetrical. The top of the structure has a 40 m decoration framework with size of the elements similar to those of the typical floor. The floor slabs adopt steel
concrete composite slabs (Fig. 8.34).

8.1.5.3.1 Building fundamental periods The building fundamental vibration modes are shown in Fig. 8.35. The first mode (T1 5 6.33 seconds) is flexural along the Y direction. The second mode (T2 5 6.26 seconds) is flexural along the X direction. The third one (T3 5 3.44 seconds) is torsion.

8.1.5.3.2 Damping strategy utilized For this project, a series of 73 viscous dampers were utilized. This included 59 dampers located in the inside tube from the 44th to the 57th story and 14 dampers located in the outside tube from the 46th to the 57th story (Figs. 8.36 and 8.37). The primary goal was to improve the service life behavior in terms of comfort under wind loads. The secondary goal was to improve the structural performance during seismic activity.

581

582

CHAPTER 8 Case studies of tall buildings

FIGURE 8.34 Beijing Yintai Center: frame support and standard layout structural plan.

FIGURE 8.35 Beijing Yintai Center: fundamental frequencies and mode shapes.

8.1.5.3.3 Additional damping provided by the damping system Although the dampers are set to provide beneficial nonlinear performance, the additional damping ratio for the first mode is approximately equivalent to 3.2%

8.1 Distributed Damping Systems Case Studies

FIGURE 8.36 Beijing Yintai Center: damper layout on floor plan.

of critical for wind in a 10-year return period. For earthquake response, the dampers provide an additional 2.2% of critical damping.

8.1.5.3.4 Building cost versus damping cost In structural design the spirit of energy conservation, environmental protection, and, economic and security integration principles lead to change the steel-section area along the height of the floors. Moreover, for the floor span two main types of steel beams are used. To reduce wind-induced vibration a certain amount of dampers installed directly can achieve the same or better effect as they strengthen the structure. Moreover, the cost of the viscous dampers makes a very economical scheme compared to increasing the size of the structural system.

583

584

CHAPTER 8 Case studies of tall buildings

FIGURE 8.37 Beijing Yintai Center: damper layout on elevation.

8.1.5.3.5 Building code Several buildings codes were utilized, such as: • • •

China code for seismic design of buildings: GB50011 (2001) Technical specification for steel structure of tall buildings: JGJ99 (1998) Guidelines for the evaluation of the response of occupants of fixed structures, especially buildings and offshore structures, to low-frequency horizontal motion (0.063
1.0 Hz): ISO 6897 (1984)

According to the referenced code for seismic design of buildings (GB5011, 2001), the maximum acceleration under frequent earthquakes, with 63% of the 100-year transcendental probability, is 85 cm/s2. Under DEs with 10% transcendental probability, the maximum acceleration is 255 cm/s2. Under unusual earthquakes with 3% transcendental probability, the maximum acceleration is 410 cm/s2.

8.1 Distributed Damping Systems Case Studies

The fundamental wind pressure in 50 years of design benchmark period is 0.5 kN/m2. Without additional damping, the maximum acceleration at the top floor exceeded the acceptability criteria set forth in JGJ99 (1998) and ISO 6897 (1984) that is equal to 0.20 m/s2 under 10-year return period wind load.

8.1.5.3.6 Peer-reviewed project The peer-review process was carried out by John A. Martin & Associates Inc.

8.1.5.4 Damping overview 8.1.5.4.1 Damping strategy Steel structures alone have very low damping ratio especially for high-rise buildings. Attaching viscous dampers to this building improved the overall damping and reduced the response during wind loading and earthquake events. The primary goal was to improve the occupant comfort under wind loads and the secondary goal was to improve the structural performance during earthquakes. Nonlinear viscous dampers were utilized in diagonal brace at locations where relative displacement between two points is predicted and greater (Fig. 8.38).

8.1.5.4.2 Damping type The typical damper utilized with base plate is shown in Fig. 8.39: 1. Piston rod; 2. piston head; 3. seal/seal bearings;’ 4. fluid; 5. cylinder; 6. end cap; 7. base plate; 8. end clevis; 9. sphrical bearing 10 bellows. The damper has the following main properties: • • • • • •

Type of damper: Nonlinear viscous damper Design life: 50 years Maximum damper force and stroke: F 5 1200 kN and D 5 6 100 mm Nonlinear constitutive law: F 5 Cvα, with α 5 0.4 and v 5 velocity (m/s) Damping constant: C 5 2000 kN(s/m)α and C 5 1500 kN(s/m)α Main vibration period: 3 seconds , T , 6 seconds

8.1.5.4.3 Structural and damping design Code and guidelines. The major reference codes utilized in the design are the following: • •

Technical specification for steel structure of tall buildings: JGJ99 (1998) Code for seismic design of buildings: GB50011 (2001)

Analysis modeling and software. Two types of analysis software were used. PKPM was used for designing the structure in China. However, this software at that time had no ability to model viscous dampers and analyze structure with dampers. Therefore, ETABS was used to analyze the building with viscous dampers. The damping devices were modeled as a Maxwell model (Chapter 4)

585

586

CHAPTER 8 Case studies of tall buildings

FIGURE 8.38 Beijing Yintai Center: viscous damper detail photograph.

FIGURE 8.39 Beijing Yintai Center: viscous damper composition diagram.

8.1 Distributed Damping Systems Case Studies

whereby a spring component is modeled in series with a dashpot (damper) component. Some of the analysis performed by ETABS were response spectrum and time history. Design principles. The basic principle was to achieve the comfort performance under wind with respect to adding damping. The structural response was calculated for both undamped and damped structures. Under wind load, the damped structure acceleration of the top floor was less than 0.20 m/s2. The acceleration time history of the 54th floor was compared for the case with and without dampers (Fig. 8.40). The performance (story drift and story shear) of the damped structure was superior to the undamped one (Fig. 8.41) (Tables 8.10
8.13).

FIGURE 8.40 Beijing Yintai Center: comparison of acceleration time history of the 54th floor with and without dampers.

FIGURE 8.41 Beijing Yintai Center: comparison of story drift and story shear.

587

588

CHAPTER 8 Case studies of tall buildings

Table 8.10 Beijing Yintai Center: Experimental Test Case Plan Case

Frequency (Hz)

Amplification (mm)

Maximum Velocity (mm/s)

Cycle Numbers

1 2 3 4 5 6 7 8

0.15 0.1 0.3 0.5 1.0 1.0 1.2 1.5

10 100 40 15 15 10 10 10

9.42 62.83 75.40 47.12 94.25 62.83 75.40 87.96

60 10 5 5 5 5 5 5

Table 8.11 Beijing Yintai Center: Acceleration of Stories Under Wind Load Acceleration X Direction (m/s2)

Acceleration Y Direction (m/s2)

Story

Without Damper

With Damper

Without Damper

With Damper

55 54 53 52 51

0.236 0.233 0.229 0.225 0.222

0.200 0.197 0.194 0.192 0.189

0.231 0.225 0.219 0.213 0.208

0.204 0.200 0.198 0.196 0.193

Table 8.12 Beijing Yintai Center: Structural Effect Under Frequent Earthquake

Basement shear (kN) Max story drift (49th story) Top floor Acceleration (m/s2)

With Damper

Without Damper

Earthquake Wave

X

Y

X

Y

Yts1 Yts4 Yts1 Yts4 Yts1 Yts4

10879.88 12743.91 1/803 1/444 0.5706 0.6898

10546.79 12194.92 1/807 1/456 0.5699 0.6948

11943.88 14099.83 1/738 1/424 0.5895 0.795

11622.01 13675.28 1/766 1/435 0.5877 0.7857

Design phase considerations. The preliminary design adopted linear dampers. More than 100 linear dampers were used to satisfy the code. However, it is well known that linear dampers do not provide the same level of efficiency as nonlinear dampers. Nonlinear dampers are able to dissipate more energy having the same conditions as linear dampers.

8.1 Distributed Damping Systems Case Studies

Table 8.13 Beijing Yintai Center: Structural Effect Under Rare Earthquake

Basement shear (kN) Maximum story drift (49th story) Top floor acceleration (m/s2)

With Damper

Without Damper

Earthquake Wave

X

Y

X

Y

Yts1 Yts4 Yts1 Yts4 Yts1 Yts4

53550.86 62214.55 1/157 1/95 3.36 3.76

51673.01 58896.25 1/158 1/97 3.36 3.76

54579.04 64483.69 1/156 1/93 3.40 3.82

52655.95 60900.73 1/155 1/94 3.40 3.80

8.1.5.4.4 Architectural integration strategy The connection of all dampers used diagonal brace as illustrated in Fig. 8.38. The damper locations were eventually covered from view by walls with a maintenance door.

8.1.5.4.5 MEP integration strategy There is no interaction or interference between the damping system and the MEP.

8.1.5.4.6 Elevator and other mechanical devices integration strategy There is no interaction between the damping system and other mechanical devices.

8.1.5.4.7 Experimental tests Before all the dampers were installed, a number of tests were carried out. The damper component level production tests were performed at the Taylor Devices Seismic Test Facility (United States). Each damper was tested to full force and velocity. Additionally, the building owner selected a third party laboratory and organized the designer, the constructor, and the supervisor to witness the damper performance tests. The test cases were established as given in (Table 8.10). The test results showed that the damper met or exceeded all the requirements of the test plan. There was no evidence of binding, yielding, or permanent deformation in any part of the damper (Fig. 8.42).

8.1.5.4.8 Monitoring system There is no monitoring system in the building.

8.1.5.4.9 Maintenance strategy Taylor Devices, Inc. provided a 35-year warranty. Taylor Devices’ FVDs are designed to be completely maintenance-free for the life of the dampers. All

589

590

CHAPTER 8 Case studies of tall buildings

FIGURE 8.42 Beijing Yintai Center: test results of damper force match the theory diagram curve.

surfaces on the damper are coated for corrosion protection. No periodic maintenance, inspection, or spare parts are required, desired, or recommended.

8.1.5.5 Comparison of different design strategies 8.1.5.5.1 Structural options considered It was very difficult to ensure concrete structure’s earthquake-resistant performance. For this reason, the entire framing was made in steel. For example, the lower transfer floor was designed, but the stiffness was too big compare to the steel superstructure.

8.1.5.5.2 Damping solution considered BRB system had been considered.

8.1.5.5.3 Cost
benefit analysis Damping cost. The cost of 73 Taylor Devices viscous dampers including all of shipping, customs duty, Chinese tax, braces, and installation fee is f323,000 (approximately USD 52,000) per damper location (Fig. 8.43).

8.1 Distributed Damping Systems Case Studies

FIGURE 8.43 Beijing Yintai Center: viscous damper in the test machine of the third part test lab.

8.1.5.6 Lesson learned and recommendations 8.1.5.6.1 Difficulties in the design The project was initially designed 10 years before. In China, at that time only a few designers and owners could accept the idea of adding viscous damping devices (VDDs) to the building. It was not easy to convince the designer and owner that the structure would achieve good performance inexpensively by adding viscous dampers.

8.1.5.6.2 Design innovative solutions Yintai Centre was the first high-rise structure with viscous dampers in China. Achieving the comfort level under wind load by adding dampers was indeed an innovation at that time. Compared to linear damper, the nonlinear damper could achieve the same effect with less quantities. It was also a big innovative solution.

591

592

CHAPTER 8 Case studies of tall buildings

FIGURE 8.44 San Diego Central Courthouse: building overview. © Skidmore, Owings & Merrill LLP | Bruce Damonte, 2018. All rights reserved.

8.1.5.6.3 Possible improvements One possible improvement could be the utilization of toggle brace or scissor brace configuration that could eventually reduce the number of damper utilized.

8.1.6 SAN DIEGO CENTRAL COURTHOUSE, SAN DIEGO, CALIFORNIA, UNITED STATES 8.1.6.1 Project data The major building data (Fig. 8.44) are summarized as follows: • • • • • • • • • • •

Year of completion: 2017 Owner / General Contractor: Superior Court of California, San Diego Central Courthouse, Judicial Council of California / Rudolph & Sletten Architectural designer: Skidmore, Owings & Merrill LLP Structural designer: Skidmore, Owings & Merrill LLP Damping device supplier and testing: Taylor Devices, Inc. Wind tunnel laboratory: RWDI Wind Tunnel Modeling Height: 118.57 m to the roof parapet Interstory height: 4.88 m Stories: 24 above grade and 2 below grade Gross area of the tower: 65,403 m2 Building function: Courthouse

8.1 Distributed Damping Systems Case Studies

•

•

Structural material Steel superstructure gravity and lateral special moment frame: ASTM A992 and A572 Gr. 50 WF and Plates Typical SMF cruciform columns W33 and box columns 33 3 33 and W24/ W30 and W36 beams VDDs: 330 kip (1468 kN) and 440 kip (1957 kN) with HSS 10 (tube steel) extender braces Floor plan area: 6,574 m2 at level 1 with 98.67 m 3 66.63 m dimensions

8.1.6.2 Introduction/History The state-of-the-art new San Diego Central Courthouse facility features 71 courtrooms comprising approximately 65,403 m2 of administrative, clerk, jury, security operations, holding, and building support spaces (Fig. 8.44). Designed for the Judicial Council of California (JCC), the Superior Court facility is located in the City of San Diego, California—a region of high seismicity situated in close proximity to active and potentially active downtown San Diego earthquake fault zones. The 118.57 m tall, 24-story above grade and two level below grade superstructure was designed to meet “enhanced” seismic performance objectives according to California Trial Court Facility Standards (CTCFS, 2011).

8.1.6.3 Structural system The superstructure lateral force resisting system features two-way ductile steel special moment frames (SMFs) with reduced beam section (RBS) connections. The SMFs are located along the perimeter in the longitudinal direction and along all gridlines in the transverse direction (Figs. 8.45 and 8.46). Columns are located typically on approximately 11.58 m 3 10.06 m, and 5.49 m 3 10.06 m modules. The typical story height of 4.88 m accommodates the typical courtroom floors with high ceiling heights. Typical SMF consists of W24 and W36 beams as well as 2-W30 cruciform columns and box columns ranging from 0.51 to 0.91 m dimension.

8.1.6.3.1 Building fundamental periods The building fundamental vibration modes are shown in Fig. 8.47: first mode (T1 5 5.22 seconds) is mainly flexural along the longitudinal direction, second mode (T2 5 4.73 seconds) is mainly flexural along the transverse direction, and third mode (T3 5 4.44 seconds) is mainly torsional.

8.1.6.3.2 Damping strategy utilized The design incorporates distributed supplemental damping with 106 VDDs with extended steel-braced elements in the slender two-bay transverse direction (see Figs. 8.45 and 8.46). Final architectural coordination did not permit locating VDD in the SMF longitudinal direction. In the transverse direction, VDDs are configured in opposing tension-compression two-bay diagonal configuration and are located at the center and extreme end gridlines (lines 1, 4, and 9) to

593

594

CHAPTER 8 Case studies of tall buildings

FIGURE 8.45 San Diego Central Courthouse: typical upper level framing plan with red VDD lines.

FIGURE 8.46 San Diego Central Courthouse: typical transverse frame elevation with red VDD lines.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.47 San Diego Central Courthouse: fundamental frequencies and mode shapes.

effectively control torsion under peak seismic demands. The VDDs are distributed along the height of the structure from level 6 to roof level with typically 6 but not less than 4 at each level (Fig. 8.46).

8.1.6.3.3 Additional damping provided by the damping system To quantify the additional damping provided by the damping devices under wind loads, a static wind load was assigned at the rigid diaphragm for each level of the structure. Then, damping ratios of the structure were calculated for each mode in order to estimate the reduction in the resonant or dynamic portion of wind load due to increased damping ratios. The inherent damping ratio of the structure for wind analysis was assumed as 1.5%. The additional damping ratio of each mode due to linear VDD was calculated by dividing the ratio of energy dissipated in one cycle ED to the total available potential energy ESo by 4π as shown in Eq. 3.60 (Fig. 3.10; Chapter 3). Therefore, the resulting additional damping ratios due to VDDs for the first 10 modes of the structure are given in Table 8.14. It is noted that because the damping devices were placed along the transverse direction only, the additional damping ratios for modes in the longitudinal direction were negligible. In the primary modes for the transverse direction and torsion, due to the utilization of VDDs, the damping ratios of the structure are 11.24% and 16.43%, respectively. In general, higher values of additional damping ratios were noted for higher modes. Therefore the resonant portion of the wind-induced load is reduced significantly by the damping devices. The VDD utilized has linear and nonlinear properties depending on the velocity values. The damper device’s linear and nonlinear properties provided by device manufacturer during the design development are shown in Fig. 8.48. The C and α coefficients are, respectively, 27143 kN-s/m and 1.0 for velocity less than or equal to 0.015 m/s and 21015 kN-s/m and 0.5 for velocity greater than 0.015 m/s. Under wind load, the velocities in the dampers are smaller than

595

596

CHAPTER 8 Case studies of tall buildings

Table 8.14 San Diego Central Courthouse: Predicted Damping Ratio With VDDs (C 5 27,143 kN-s/m) Mode

Period

Type

ζ (%)

1 2 3 4 5 6 7 8 9 10

5.22 4.73 4.44 1.91 1.76 1.67 1.36 1.11 1.03 1.01

Longitudinal Transverse Torsion Longitudinal Transverse torsion Transverse torsion Vertical Longitudinal Torsion Transverse

0.09 11.24 16.43 0.49 41.64 45.28 0.00 0.37 91.89 74.60

FIGURE 8.48 San Diego Central Courthouse: nominal force
velocity relation of VDD (Taylor Devices).

0.015 m/s. Therefore, for the purpose of the wind study, the VDD elements were modeled with C and α values of 27143 kN-s/m and 1.0, respectively. From all of the three studies, the model with damping properties of C 5 27143 kN-s/m (the lower bound damping coefficient) and α 5 1 provides additional damping ratios (for the fundamental transversal mode) of 11.2%, 10.9%, and 15.6%, respectively per each method. In the final analytical model for wind analysis, conservatively 6% damping ratio was used in the transverse direction (transverse and torsional mode) and 1.5% damping ratio in longitudinal direction. Therefore, it was specified that the VDD prototype testing under low amplitude and velocity (per ASCE

8.1 Distributed Damping Systems Case Studies

7-05 (ASCE, 2005)) achieve damping levels of 6% or greater. Three studies were conducted to evaluate the damping ratio of structure with VDDs with specified damper properties under wind loads: (1) the modal properties, (2) the dissipated energy in the system under free vibration, or (3) the decay of motion in free vibration.

8.1.6.3.4 Building cost versus damping cost Information is not available.

8.1.6.3.5 Building code The 118.6 m tall, 24-story above grade and two level below grade superstructure was designed to meet “enhanced” seismic performance objectives of the CTCFS (2011), Section 12D Criteria for Rare Loads based on the CBC (2010) and ASCE 7-05 (ASCE, 2005) using site-specific seismic hazard and ground motion design criteria. Prescriptive building code minimum requirements: CBC (2010), ASCE 7-05 (ASCE, 2005), AISC 360-05 (AISC, 2005), AISC 341-05 (2010), AISC 358-05 (ASCE, 2005). Nonlinear inelastic performance based: ASCE 41-06 (ASCE, 2007)/FEMA 356 (ASCE, 2000).

8.1.6.3.6 Peer-reviewed project JCC project design and ASCE 7-05 (ASCE, 2005) required peer review that which was provided by John A. Martin & Associates, Inc. with Michael Constantinou, Gregory Deierlein, and Norman Abrahamson.

8.1.6.4 Damping overview 8.1.6.4.1 Damping strategy Different configurations of dampers were studied along with the different distribution of dampers along the height of the building primarily for controlling the seismic drifts and accelerations. With the final layout and sizes of the VDD in the frame based on seismic design, the study for effectiveness of these VDDs in wind design was then undertaken.

8.1.6.4.2 Damping type The VDDs provide an output force in either tension or compression that is directly related to the relative velocity between the two ends of the dampers. The damper output force varies only with velocity and does not change with damper stroke position or orientation angle. The function of the dampers is primarily to absorb earthquake energy, thereby reducing or eliminating damage to the building when an earthquake occurs. The San Diego Central Courthouse VDD properties had a nonlinear relation to the velocity except for low velocities

597

598

CHAPTER 8 Case studies of tall buildings

(less than 0.015 m/s), applicable under wind load conditions where the relation was linear. The device as shown in Figs. 8.49
8.52, in its diagonal configuration, connects the moment frame beam
column joint at one end and the extender brace at the other end. The other end of the extender brace connects to the diagonally opposite moment frame beam
column joint. This unit uses inert silicone fluid as the operating fluid medium which complies with US Federal Standard VV-D-1078.

FIGURE 8.49 San Diego Central Courthouse: typical damper and extended brace diagonal frame configuration.

FIGURE 8.50 San Diego Central Courthouse: (A) VDD detail in-place and (B) VDD detail drawings.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.51 San Diego Central Courthouse: (A) VDD typical, (B) SMF RBS cruciform column, and (C) SMF RBS box column in-place.

FIGURE 8.52 San Diego Central Courthouse: (A) typical VDD damper connection detail at SMF and (B) typical VDD extender brace connection detail at SMF.

8.1.6.4.3 Structural and damping design Code and guidelines. The wind loading conditions were determined based on CBC (2010), ASCE 7-05 (ASCE, 2005), Chapter 6. With an occupancy category of III, the structure is designed for a wind importance factor of 1.15. The surface roughness category was determined to be B and the exposure category was C. The wind tunnel testing report RWDI (2011) provided the overall wind-induced loads for each wind direction. These loads were based on the design wind speed, assuming this wind speed applies equally to all directions and no allowance is made for the relative probability that the design wind speed will occur from different location. Analysis modeling and software. A three-dimensional (3D) ETABSv9.7 (CSI, 2011) analytical model was built with linear steel frame and nonlinear VDDs for seismic response history analysis and as linear VDDs for wind analysis. In the structure the VDDs are installed with the HSS extenders diagonally spanning in the SMF as shown in Figs. 8.49 and 8.50. The force in VDD is

599

600

CHAPTER 8 Case studies of tall buildings

determined by the equation F 5 Cvα, where damping coefficient (C) and damping exponent (α) are the properties of VDD elements, and the VDDs do not contribute any stiffness to the global structural frame. In the ETABS analytical model, the VDDs were assigned in the HSS extender elements as nonlinear link “damper” elements with initial stiffness, damping coefficient, and damping exponent. Design principles. The focus of this work is to identify the effectiveness of VDDs, employed primarily in the case of the new San Diego Central Courthouse to provide energy dissipation and damping under high rates of seismic dynamic loading, and to also provide a reduction of the dynamic component of motion under low amplitude wind loads. The JCC “enhanced” seismic performance objectives are intended to reduce losses beyond a code baseline “normal” seismic performance objective per Chapter 12 of the California Trial Court Facilities Standard (Fig. 8.53; CTCFS, 2011). Where an “enhanced” objective is directed by the JCC (2006), it is expected that the structure may experience minor to moderate inelastic demands allowing economical repairs, while damage to nonstructural components, contents, as well as potential loss of functionality and business interruption are expected to be significantly reduced with respect to a “normal” design based on minimum building code values. “Enhanced” seismic performance objectives are effectively achieved through the use of VDD technologies combined with steel moment frames in tall building

FIGURE 8.53 San Diego Central Courthouse: expected performance level as related to occupancy category (OC) and level of ground motion (FEMA P-750 (NEHRP, 2009)).

8.1 Distributed Damping Systems Case Studies

structural systems. The nonlinear VDDs are utilized to economically control damage in tall buildings when subjected to moderate to major ground shaking earthquake by significantly limiting building story drifts, floor accelerations, and base shear demands. The JCC (2006) designated an “enhanced” seismic performance objective for the design of the San Diego Central Courthouse project. In collaboration with the peer and design review panel, a comprehensive twostage seismic design criteria was devised meeting the prescriptive requirements of ASCE 7-05 (ASCE, 2005) structural modeling, analysis, and acceptance criteria. Stage 1 included modeling of linear elastic moment frame elements with nonlinear VDDs using bidirectional response history analysis based on DE and maximum considered earthquake (MCE) site-specific ground motions. Stage 2 extended the Stage 1 modeling and analysis to model the nonlinear plastic hinge properties of the beam “RBS” connections and capture inelastic demands in compliance with ASCE 41-06 (ASCE, 2006) acceptance criteria. Bounded analyses were performed based on upper and lower bound (615%) VDD properties. Additionally, effective damping of the nonlinear VDDs under resonant-induced wind loading was evaluated using modal energy methods. Preliminary damper testing was performed to assess damper device effectiveness at low velocities. As part of the design development and construction document phases, a steel moment connection qualification testing program was successfully undertaken to provide qualification acceptance of large-scale steel box column (0.61 m 3 0.91 m 3 0.05 m) and beam “RBS” (W36 3 302) sections for the project. During early design development phases, seismic risk and life cycle assessments were completed to inform decision-making in the selection of alternative structural systems leading to cost-effective enhanced seismic performance objectives as discussed in Section 8.1.6.5.3. Design phase considerations. In agreement with the peer and design review panel, a two-stage structural modeling, analysis, and design approach was adopted in conformance with the prescriptive provisions of CBC/ASCE 7 requirements. In Stage 1, CBC (2010) and ASCE 7-05 (ASCE, 2005) code compliance was demonstrated using linear elastic analysis of the SMF with nonlinear VDD designed to meet basic code provisions of ASCE 7-05 (ASCE, 2005) emphasizing detailed requirements of chapter 18, “Seismic Design Requirements for Structures With Damping Systems” and the damper and seismic force resisting system (SFRS) frames using a 3D ETABS (CSI, 2011) structural model. This linear analysis assumes that the SMF demand/capacity (D/C) ratios from the unreduced seismic load combinations are less than or equal to 1.5. In accordance with ASCE 7-05 (ASCE, 2005) Section 18.3 provisions, which stipulates that when D/C ratios in the SFRS frame elements are greater than 1.5, inelastic behavior shall be modeled using a Stage 2 nonlinear analysis procedure at DE level demands. In Stage 2, inelastic hinges were modeled in the SMF beams using nonlinear link elements of ETABS (CSI, 2011). Inelastic hinge rotation demands were checked using ASCE 41-06 (ASCE, 2007) and FEMA 356 (FEMA, 2000) acceptance criteria. SMF columns were designed to remain elastic under the DE combined with expected

601

602

CHAPTER 8 Case studies of tall buildings

gravity loading in both the Stage 1 and Stage 2 analyses, meeting the intent of biaxial RBS moment connection prequalification of AISC 341-05 (AISC, 2005). Table 8.15 summarizes seismic design and acceptance criteria for both the Stage 1 and Stage 2 structural modeling, including seismic force reduction factors in each frame direction used in the linear elastic analysis as well as beam hinge rotational limits used as design acceptance criterion for the inelastic nonlinear analysis. Building story drift was limited to 1.5% per ASCE 7-05 (ASCE, 2005) at the DE for the occupancy category III—defined courthouse facility. Stage 1 modeling. In addition to seismic demands using response history analysis, a load case was defined to account for accidental torsion moment of 5% of the longest horizontal dimension of each story times the equivalent lateral force from ASCE 7-05 (ASCE, 2005) Section 12.8. The direct analysis method was used to design the superstructure frame and diaphragm in accordance with AISC 360-05 (ASCE, 2005), Appendix 7. The code commentary reduced stiffness was applied to the models, used in analyses of strength limit states, and not for the models used for serviceability checks, such as story drift and ASCE 7-05 (ASCE, 2005) Section 12.3 structural irregularities. At the below grade levels, the SMF was designed as part of a dual system with perimeter reinforced concrete (RC) shear walls (ASCE 7-05 (ASCE, 2005) Table 12.2-1; R 5 7, Ω 0 5 2.5, Cd 5 5.5), where the moment frame without the basement walls is required to be capable of resisting 25% of the structure base shear. There are two types of SMF beam to column moment connections in the SFRS—at the RBS beam (Fig. 8.54) and at the beam moment connection just outside of the damper brace extender, that consists of the built-up gusset plate connections. Each was modeled to take into account the change in stiffness due to variation in cross-sectional second moment of area. The moment connection at the damper and extender brace was designed using capacity concepts to allow for a hinge formation at a distance db/2 from the edge of the damper and extender brace gusset plate, where db is the depth of the moment frame beam at the connection. The VDDs were idealized in the ETABS (CSI, 2011) model using nonlinear link “damper” elements. The initial stiffness, damping constant (C), and damping exponent (α) were input to the analysis model. The dampers were modeled with lower bound and upper bound damping properties ( 6 15% C coefficient) and nominal specified strength and stiffness in the analysis and design. For this project, VDD nonlinear properties with C 5 21015 kN-s/m and α 5 0.5 were used. Vertical seismic mass was included in fast nonlinear analysis modeling together with 725 modes to properly capture the response of the “damper” nonlinear link elements. In Stage 2, the Stage 1 ETABS (CSI, 2000) model was modified to account for the formation of plastic hinges in the SMF beams. Each RBS beam was divided into five segments. Segments 1, 3, and 5 are linear nonprismatic and represent the variation in the cross-section of the beam. Segments 2 and 4 contain 1-inch (2.5 cm) long linear prismatic beam elements which have zero flexural stiffness in parallel with nonlinear links that have the nonlinear moment
rotation relationship and properties.

Table 8.15 San Diego Central Courthouse: Seismic Design Criteria Summary Summary Seismic Design Criteria

Seismic Force Reduction Factors

Minimum Strength Design Base Shear (DE)

Acceptance Criteria

Structural Modeling

Structural Analysis

Structural System

Code Standards

State 1: ETABS (3D)

Linear Elastic MF Nonlinear VDD Site-specific Bidirectional response history (RHA) ASCE 7-05 (ASCE, 2005) Ch. 16.1 and 16.3 (ζ 5 5%) Design earthquake Maximum considered earthquake MCE

EW direction MF 1 VDD

CBC (2010) ASCE 7-05 (ASCE, 2005) Ch. 12, 16, and 18 AISC 34105 (ASCE, 2005) Occupancy category III

R/ (Cd Ω0) SFRS MF DE: R58 I 5 1.25 Cd 5 5.5 Ω0 5 3

0.75Vmin per ASCE 705 (ASCE, 2005) Sec. 18.2.2.1 Vmin per ASCE 705 (ASCE, 2005) Sec. 12.8.1

DE Δx 5 Δxe Δx # 1.5%R/Cd

DE: CBC (2010) Ch. 1605.2.1

NS direction MF

CBC (2010) ASCE 7-05 (ASCE, 2005) Ch. 12 and 16.1 AISC 34105 (ASCE, 2005) Occupancy category III

I/R SFRS MF DE: R58 I 5 1.25

0.85Vmin per ASCE 705 (ASCE, 2005) Ch. 16.1 Vmin per ASCE 705 (ASCE, 2005) Ch. 12.8.1

n/a

DE: CBC (2010) Ch. 1605.2.1

Drift Limit

Load Cases

Performance Objective

DE

MCE

SFRS MF ASCE 7-05 (ASCE, 2005) Sec. 18.4 RHA V $ 0.75Vmin Preliminary DS frame (VDD) D/C # 1.0 Nominal ϕ 5 1.0 AISC 360-05 (ASCE, 2005) AISC 341-05 (ASCE, 2005) AISC 358-05 (ASCE, 2005) SFRS MF ASCE 7-05 (ASCE, 2005) Ch. 16.1 RHA V $ 0.85Vmin AISC 360-05 (ASCE, 2005) AISC 341-05 (ASCE, 2005) AISC 358-05 (ASCE, 2005)

Preliminary ASCE 7-05 (ASCE, 2005) Ch. 18 R51 VDD stroke, force, and velocity

DE: 475 ARP “Life safety” “Damage control” MCE: 2475 ARP “Collapse prevention”

n/a

DE: 475 ARP “Life safety” “Damage control” MCE: 2475 ARP “Collapse prevention”

(Continued)

Table 8.15 San Diego Central Courthouse: Seismic Design Criteria Summary Continued Summary Seismic Design Criteria

Structural Modeling

Structural Analysis

Structural System

Code Standards

State 2: ETABS (3D)

Nonlinear inelastic MF Nonlinear VDD Site-specific bidirectional response history (RHA) ASCE 7-05 (ASCE, 2005) Ch. 16.1 and 16.3 (ζ 5 2.5%) Design earthquake Maximum considered earthquake MCE

EW direction NS direction MF 1 VDD

ASCE 4106 (ASCE, 2006) ASCE 7-05 (ASCE, 2005) Ch. 16.2 and 18 NL RHA AISC 34105 (ASCE, 2005)

Seismic Force Reduction Factors

Minimum Strength Design Base Shear (DE)

R51

n/a

Acceptance Criteria

Drift Limit

DE Δx 5 Δxe Δx # 1.5%R/Cd (EW Dir. Only) DE Δx # 1.25 1.5% R/Cd (NS direction only)

Load Cases

DE: ASCE 41-06 (ASCE, 2006) Ch. 3.4.3 NL RHA

DE

MCE

NL RHA ASCE 7-05 (ASCE, 2005) Ch. 18.3 SFRS MF Inelastic beams ASCE 41-06 (ASCE, 2006) Ch. 3 and 5 θ , 6θy Elastic columns ϕ 5 1.0 Ry 5 1.1 D/C # 1.0

SFRS MF Inelastic beams ASCE 4106 (ASCE, 2006) Ch. 3 and 5 θ , 8θy ASCE 7-05 (ASCE, 2005) Ch. 18 R51 VDD stroke, force, and velocity

Performance Objective

DE: 475 ARP “Life safety” “Damage control” MCE: 2475 ARP “Collapse prevention”

8.1 Distributed Damping Systems Case Studies

FIGURE 8.54 San Diego Central Courthouse: typical SMF RBS connection at cruciform column.

Because the building has 1028 moment frame beams and 106 VDDs, the minimum number of FNA Ritz vector modes required to capture the behavior of the nonlinear link elements alone was 1028 3 2 1 106 5 2162 modes, such that a total of 3000 modes were used in the Stage 2 NLRHAs. Due to limitations in the generated output file size, the gravity framing and floors present in the Stage 1 model had to be removed and their mass and loads were lumped at joints at respective locations on every level of the Stage 2 model. Before performing NLRHA runs, the Stage 2 model was verified to have the same gravity loads, mass, mode shapes, and periods as the Stage 1 model. Additionally, the limited size of the output file required that each response history analysis was executed as a separate model.

8.1.6.4.4 Architectural integration strategy Although it was demonstrated that providing VDD frames in both orthogonal directions would have been advantageous, architecturally the VDDs could only be located in the transverse direction integrated with the partition walls on Gridlines 1, 4, and 9 without compromising the interior layout within the building. The VDDs are distributed along the height of the structure from level 6 to roof level with typically 6 but not less than 4 at each level. The final design of the structure

605

606

CHAPTER 8 Case studies of tall buildings

included SMFs in both directions with VDD as the “enhanced” seismic system in one direction only. As further noted in Section 8.1.6.5.2, which describes alternate damping configurations considered, it should be emphasized that given that VDDs do not need to be continuous to lower levels and foundations to be effective, the architectural integration of VDDs has significant advantages over conventional steel-braced frames or RC shear walls systems which often need to avoid vertical discontinuities of the SFRS and extend continuously to lower levels and foundations. In fact, for the San Diego Central Courthouse, the VDDs stopped at level 6 primarily due to significant architectural program conflicts at lower levels including large jury assembly, below grade parking, security and access issues. This was accomplished without sacrificing structural performance and economical use of VDDs.

8.1.6.4.5 Sustainability and MEP integration strategy Among the strategies for reducing energy use include radiant cooling in public corridors and lobby floors and the use of district chilled water. An optimized building fac¸ade was designed to maximize daylight but reduce heat gain through a combination of a large overhang, frit patterns in the glazing elements of the curtain wall, and high-performance glass. Other highlights of the sustainability strategy included: • • • • •

Certified LEED v3 Silver compliance. The landscape incorporates drought-tolerant native plants reducing potable water used for irrigation by 71%. Indoor potable water reduction of 37% is achieved through low flow and efficient fixtures. The building achieves a 30% reduction of energy compared with ASHRAE 90.1 2007 resulting in 18% savings in energy cost. The project occupies a previously developed site and promotes the use of alternative mode of commuting.

8.1.6.4.6 Elevator and other mechanical devices integration strategy Information is not available.

8.1.6.4.7 Prototype Damping Device Testing Construction phase VDD prototype tests conducted in January 2015 at Taylor Devices, Inc. test facility in North Tonawanda, NY, included four prototype damper devices (two each of 1468 and 1957 kN unit (Figs. 8.55 and 8.56) tested in conformance with ASCE 7-05 (ASCE, 2005), Chapter 18 requirements. Velocity performance (seismic), cycle (wind), and friction tests were conducted. Velocity performance testing met hysteresis loop acceptance criteria within 6 15% of the average of 5 cycles. Cyclic (wind) testing (each with 2000 cycles) demonstrated devices had essentially linear viscous behavior under the conditions of wind (amplitude of 4.06
4.32 mm and velocity of 5.08 mm/s) confirming

8.1 Distributed Damping Systems Case Studies

FIGURE 8.55 San Diego Central Courthouse: view of 440 kip (1957 kN) SN001 damper in large tester.

FIGURE 8.56 San Diego Central Courthouse: view of 330 kip (1468 kN) SN001 damper in small tester.

607

608

CHAPTER 8 Case studies of tall buildings

effective damping levels ranging from 14.1% to 20.0%. These are consistent with and exceed the assumed analytical predictions of 11.2% and 16.4% used in the design of the courthouse superstructure. The normal operating force developed by the VDD unit over the design range of velocity is required to be within the upper and lower bounds of 15% in both directions of travel. The force
velocity displacement and damping properties used for the design of the damping system are to be based on the prototype tests in accordance with ASCE 7-05 (ASCE, 2005), Section 18.9. Thus the prototype testing consisted of velocity performance tests, wind cyclic tests, and friction tests were performed separately on two full size damping devices for each of the 1468 and 1957 kN dampers. The tests were intended to determine the devices’ usable life and resistance to damage, so as to verify that the devices will perform appropriately when installed in the structure. These tests also provide results that validate the damper properties used in the mathematical model. The velocity performance testing primarily confirmed the damper properties used for the seismic design. Each unit was sinusoidally tested in accordance with ASCE 7-05 (ASCE, 2005), Chapter 18, Section 18.9.1.2 (2). Since the damping devices are expected to be subjected to wind-induced forces and displacements, low amplitude cyclic test was performed in accordance with ASCE 7-05 (ASCE, 2005), Chapter 18, Section 18.9.1.2 (1). The amplitude for the wind cyclic test was determined to be 14.32 mm, which is the component of the maximum drift in the direction of the damper, due to dynamic (resonant) loading for a 20-year return period wind. The imposed motion was harmonic with frequency of 0.2 Hz (inverse of the fundamental period of the building or 4.73 seconds), resulting in peak velocity of 5.08 mm/s. Considering a 3-hour windstorm, the number of cycles expected for the structure with fundamental period of 4.73 seconds is (3 3600/4.73 5 ) 2283 cycles. This is comparable to the 2000 continuous fully reversed cycles required by ASCE7-05 (ASCE, 2005). Each device was cycled for at least 2000 cycles at a displacement of 14.32 mm and a peak velocity of 5.08 mm/s. Example results for 1957 kN cyclic wind test are shown in Figs. 8.57 and 8.58.

8.1.6.4.8 Monitoring system The new San Diego Central Courthouse has installed strong motion instrumentation system to collect and process data from significant strong motion records (5% acceleration or greater at the ground level) that are obtained during earthquakes. The system was designed, specified, and procured during the base building design, construction document, and construction administration phases under the direction of the JCC and general contractor. Installation and commissioning support, and long-term maintenance of the system are provided by the California Strong Motion Instrumentation Program (CSMIP), in the California Geological Survey of the Department of Conservation, as part of the statewide network of instrumented buildings. Out of total 28 accelerometers, 24 were installed in the main building superstructure (19 horizontal and 5 vertical sensors) and 4 in the connecting

8.1 Distributed Damping Systems Case Studies

FIGURE 8.57 San Diego Central Courthouse: time history for 440 kip (1957 kN) SN001—wind test (2000 cycles).

FIGURE 8.58 San Diego Central Courthouse: force-displacement behaviour for 440 kip (1957 kN) SN001—wind test (2000 cycles).

cantilevered pedestrian bridge to adjacent HOJ at level 3 (3 horizontal and 1 vertical sensors). Sensors are distributed over the height of the main building starting at level B2 foundation (6 sensors), level 1 (2), level 4 (3), level 6 (2), level 12 (3), level 17 (3), level 23/Mech PH (4) and level 25/Roof (1). All accelerometers

609

610

CHAPTER 8 Case studies of tall buildings

are interconnected to centrally located computer controlled digital recording equipment (RefTek equipment by Trimble) and cabling for common start, GPS timing, and synchronization. At the high roof level, a GPS antenna and junction box is installed to provide support for system. Existing CSMIP accelerometers located in the vicinity at adjacent sites will provide a downhole sensor to record “free-field” earthquake ground motions. The data collected during an event provide valuable information to the building owner, maintenance personnel, and structural engineering design team to assess the performance of the building following moderate to major earthquake events. All data will be made available, in raw and processed form, soon after the earthquake at http://www.strongmotioncenter.org.

8.1.6.4.9 Maintenance strategy Following final construction and special inspection of the VDD installation, damping devices shall have a periodic monitoring, inspection, testing schedule, and maintenance program to ensure that the devices respond in a dependable manner throughout the design life in compliance with building code (CBC (2010)/ASCE 7-05 (ASCE, 2005) Sections 18.2.5.3, 18.9) requirements. During construction and final installation, means of access for inspection and replacement of damping devices shall be provided. The degree of inspection and testing shall reflect the established in-service history of the damping devices and the likelihood of change in properties over the design life of the devices. The maintenance program should include remodeling, repair, retrofitting, or replacement at the damping device location. The periodic inspection schedule for the VDDs at the new San Diego Central Courthouse includes external inspection of devices following completion of construction—within 2 months, 1 year, and every 5 years thereafter. Damping devices are subject to external and internal inspections following fire, flood, and earthquakes with device movements greater than 25.4 mm. Specified damping devices are constructed to be maintenance-free for a period of minimum 35 years. Removal and testing of devices after external and internal inspection shall be determined by responsible engineer. If required, pressure vessels and seals are to be tested to rated pressures and acceptance levels.

8.1.6.5 Comparison of different design strategies 8.1.6.5.1 Structural options considered To achieve an “enhanced” seismic performance objective during the design development phase, several structural system options were considered and evaluated. These options included SMF with supplemental VDDs, roof hybrid seismic isolation systems, and SOM patented pin fuse frame. NLRHA was completed using site-specific seismic hazard design criteria. Optional frame systems were evaluated using a 2D moment frame models subjected to both DE and MCE loads. The dynamic properties and responses of the enhanced seismic options were compared

8.1 Distributed Damping Systems Case Studies

to the baseline case of an SMF designed to code minimum or a “normal” seismic performance objective. It was observed that the SMF 1 VDD frames were more effective in reducing story drifts, shears, and floor accelerations for both DE and MCE demands. Due to the slender shape of the structure, VDDs proved to be more effective in the transverse direction. Although it was demonstrated that providing VDD frames in both orthogonal directions would have been advantageous, architecturally the VDDs could only be located in the transverse direction integrated with the partition walls on Gridlines 1, 4, and 9 without compromising the interior layout within the building. The final design of the structure included SMFs in both directions with VDD as “enhanced” seismic system only in one direction.

8.1.6.5.2 Damping solution considered For a conventional 24-story SMF building, the maximum drifts typically occur at middle third of the structure. Keeping this in view, four different damper distribution configurations were studied along the height of the building which included: (1) dampers located only at the main roof level; (2) dampers distributed from level 6 to the roof level; (3) dampers distributed from level 1 to roof level; and (4) dampers distributed from level 4 to level 16 (Fig. 8.59). For these preliminary studies, a constant damping of C 5 7711 kNs/m was assumed for all the frames

FIGURE 8.59 San Diego Central Courthouse: lateral frame distribution scenarios over height.

611

CHAPTER 8 Case studies of tall buildings

with VDDs, and the corresponding results obtained from NLRHA were compared with the baseline SMF case. The comparative story drift and floor accelerations at the DE of the lateral frames for different damper distributions are shown in Figs. 8.60 and 8.61. The frame with dampers at the roof had similar performance as the baseline case scenario. The remaining three distributions of damper scenarios performed similar and resulted in significant reduction in story drifts and floor accelerations. It is noted that story drift ratios and floor accelerations are the two most commonly used indicators of damage to deformation-sensitive and acceleration-sensitive nonstructural components which contribute to primary costs of repair after an earthquake. Considering these 2D frame analysis results, the scenario with dampers distributed from levels 4 to 16 would have made more economic sense for the project.

122

107

91

76 Elevation (m)

612

61 Roof damper 46

L04–L16 damper L06–roof damper

30

L01–roof damper Base

15

0 0

0.01

0.02

0.03

0.04

0.05

Inter–story drift ratio

FIGURE 8.60 San Diego Central Courthouse: story drift ratios of different damper configurations.

8.1 Distributed Damping Systems Case Studies

122

107

91

Elevation (m)

76

61

46

30 Roof damper L04–L16 damper L06–roof damper

15

L01–roof damper Base

0 0

0.1

0.2

0.3

0.4

0.5

0.6

Floor acceleration (g)

FIGURE 8.61 San Diego Central Courthouse: floor acceleration of different damper configurations.

However, to achieve an economical control of peak seismic torsion demands in the transverse direction of the 3D analysis modeling, the option with dampers extending from level 6 to penthouse roof level was shown to be more effective. It is further noted that, due to the fact that there is a high roof cornice in the structure, it was preferred to distribute the dampers up to the roof level to reduce the significant seismic forces induced at the roof level. In addition, for architectural flexibility, all dampers below level 6 were eliminated to preserve large usable open space. As a result of these comparative studies, the final damper distribution chosen for this project was from level 6 to roof level. It is important to note that unlike typical braced frame systems, the dampers need not be continued to the foundation or constrained to one frame elevation, allowing additional flexibility. They can be provided at optimum locations both in plan and over the height of structure where they are determined to be most effective in the SMF system.

613

614

CHAPTER 8 Case studies of tall buildings

8.1.6.5.3 Seismic Risk Assessment and Life-Cycle Analysis During the design development phase of the new San Diego Central Courthouse, additional seismic risk assessment and life-cycle analysis were conducted to aid the JCC in the design decision-making. To evaluate the predicted seismic performance and long-term benefit with respect to additional first costs of adding VDD to the SMF lateral system, studies included a baseline “normal” seismic performance objective using a minimum code compliant SMF (Occupancy Category III), in comparison with “enhanced” seismic performance criteria for both SMF (Occupancy Category IV) and SMF 1 VDD (Occupancy Category III) systems. A standardized earthquake loss estimation methodology was utilized based on HAZUS (FEMA, 2002) which includes comprehensive sets of building damage and loss functions for the structural system, nonstructural systems sensitive to both building drift and floor accelerations, and building contents. Additionally, impacts due to business disruption and relocation expenses were assessed based on JCC damage and cost models. Simplified nonlinear capacity (pushover) curves and ground motion spectra analyses in each principle direction were used to estimate economic losses resulting in expected mean annual loss, cost–benefit ratios, and return on capital investment based on a 25-year life cycle. Results for three SMF systems are summarized in Table 8.16 which illustrates the long-term benefit of the “enhanced” SMF 1 VDD system. It is further noted that this option included SMF steel quantities based on the “normal” baseline option. In this regard, it is concluded that the VDDs are more effective and economical in controlling seismic drift demands for the long-term LCA than the additional SMF steel tonnage required for the “enhanced” SMF option which illustrated a 2 10.6% annual return on investment.

8.1.6.6 Lesson learned and recommendations 8.1.6.6.1 Difficulties in the design Information is not available. Table 8.16 San Diego Central Courthouse: Seismic Risk Assessment and LCA Structural System Option

Additional First Cost

Average Annual Loss

Average Annual Return

Business Relocation (days)

Annual Return on Investment

Normal (SMF) Enhanced (SMF) Enhanced (SMF 1 VDD)

Baseline

$804,000

Baseline

140

Baseline

$6.1m

$763,000

$42,000

125

2 10.6%

$5.5m

$346,000

$458,000

0

1 6.7%

8.1 Distributed Damping Systems Case Studies

8.1.6.6.2 Design innovative solutions Information is not available.

8.1.6.6.3 Possible improvements Possible improvements include: 1. Provide similar distribution of viscous damping devices in both longitudinal direction of building and transverse direction as provided. This was considered during design development and showed significant benefits of steel moment frame efficiency in and enhanced seismic performance. 2. Utilization of alternate toggle brace or scissor brace configuration that could increase effectiveness of damping devices and eventually reduce the number of damper devices.

8.1.7 WUHAN POLY CULTURAL PLAZA, WUHAN, CHINA 8.1.7.1 Project data The major building data (Fig. 8.62) are summarized as follows: • • • • •

•

•

•

• • • •

Year of completion: 2012 Developer: Hubei Poly Real Estate Co., Ltd (China) Contractor: General Construction Company of CCTEB Group Co., Ltd (China) Owner: Poly Real Estate Co., Ltd (China) Architectural designer: Concept design: Skidmore, Owings and Merrill (United States) Optimization design: Central South Architectural Design Institute (China) Structural designer: Concept design: Skidmore, Owings and Merrill (United States) Optimization design: Central South Architectural Design Institute (China) Damping supplier: Taylor Devices, Inc. (United States) BEIJING QITAI Shock Control and Scientific Development Co., Ltd (China) Testing laboratory Wind Tunnel Testing: Institute of Structural Wind Engineering, Civil Engineering College, Wuhan University (China) Taylor Devices (United States) Height: 211.8 m from street level Interstory height: 2.8 m (standard level) Stories: 46 above grade (main building) 1 20 above grade (vice building) 1 8 above grade (annex) 1 4 below grade Gross area of the tower: Covers an area of 12,200 m2, total surface 142,800 m2

615

616

CHAPTER 8 Case studies of tall buildings

FIGURE 8.62 Wuhan Poly Cultural Plaza: building overview.

• • •

Building function: 5A grade comprehensive office building Structural material: Type of dampers: 67DP-18900-01, 67DP-18901-01, 67DP-18902-01 Floor plan area: L-shaped plan from levels 16 to 21 having dimensions L 5 59.5 m 3 89.25 m, total floor area 3793 m2 Rectangular plan from levels 22 to 46 having dimensions L 5 59.5 m 3 25.5 m, total floor area 1517 m2

8.1 Distributed Damping Systems Case Studies

8.1.7.2 Introduction/history The Wuhan Poly Culture Plaza building, located in downtown Wuhan, is a 5A comprehensive office building. The building covers an area of 12,000 m2, with a total construction area of about 140,000 m2, and for the podium on the ground floor a total area of 44,210 m2. The overall layout of the building is L type, and the main building is connected with the auxiliary building by a connecting body. The podium relates them closely and forms a structural unit together. The main and auxiliary building structure adopts round steel tube concrete columns, steel beams, and RC core wall structure. The two buildings are connected at the 16
20 floors with a steel truss structure.

8.1.7.3 Structural system Steel frame RC core wall hybrid structure system (see Table 8.17 for structural material quantities utilised in the building).

8.1.7.3.1 Building fundamental periods The building fundamental periods as shown in Fig. 8.63 are the following: • • •

The first vibration mode: translation of Y direction with a period of 6.26 seconds The second vibration mode: translation of X direction The third vibration mode: torsional vibration

8.1.7.3.2 Damping strategy utilized Dampers were set in the weak layers, with large displacement under the earthquake action and were arranged symmetrically (Figs. 8.64
8.66).

8.1.7.3.3 Additional damping provided by the damping system Under the frequent earthquake, the additional damping ratio is 3.0% in the X direction and 2.5% in the Y direction. Table 8.17 Wuhan Poly Cultural Plaza: Structural Material Quantities The total amount of concrete (m3)

76,000

Concrete discount thickness per square meter (cm/m2)

The total amount of steel (tons)

Steel bar: 14,000

Steel consumption per square meter (kg)

Shape steel: 15,000

117 (underground) 32 (overground) Steel bar: 210 (underground) 60 (overground) Shape steel: 30 (underground) 120 (overground)

Types of dampers: 67DP-18900-01, 67DP-18901-01, 67DP-18902-01.

617

618

CHAPTER 8 Case studies of tall buildings

FIGURE 8.63 Wuhan Poly Cultural Plaza: fundamental frequencies and mode shapes.

FIGURE 8.64 Wuhan Poly Cultural Plaza: structural model and floor plans.

Under the rare earthquake, the additional damping ratio is 1.0% both in the X direction and Y direction.

8.1.7.3.4 Building cost versus damping cost Total investment of building: more than f1.1 billion (US$175 million).

8.1 Distributed Damping Systems Case Studies

FIGURE 8.65 Wuhan Poly Cultural Plaza: damper layout diagram.

FIGURE 8.66 Wuhan Poly Cultural Plaza: construction view.

8.1.7.3.5 Building code Several buildings codes were utilized and some of others were considered as reference documents. Among all, the most important are: • •

GB50011 (2010) code for seismic design of building JGJ3 (2010) technical specification for concrete structures of tall building

619

620

CHAPTER 8 Case studies of tall buildings

8.1.7.3.6 Peer-reviewed project The project passed the Hubei Province high-rise buildings seismic engineering special review; and passed the construction drawing review of the Hubei JianE Survey and Design Review & Consulting Co., Ltd.

8.1.7.3.7 Design forces A set of both natural and artificial records were utilized in the design analyses: • •

Natural records: El Centro and Taft ground motions Artificial records: one of the records in the 83 hole of the local surface soil reaction, whose transcendental probability is above 2%

The acceleration peak value is equal to 192 gal, mainly in the horizontal direction. Based on the analyses studies the following base shear forces were computed for both frequent and rare earthquakes: Base shear force of frequent earthquake: • X direction maximum: 9168 kN (no damper) and 8473 kN (with dampers) • Y direction maximum: 6161 kN (no damper) and 5525 kN (with dampers) Base shear force in rare earthquake: • X direction maximum: 97,800 kN (no damper) and 96,982 kN (with dampers) • Y direction maximum: 65,718 kN (no damper) and 63,904 kN (with dampers)

8.1.7.3.8 Expected performance Story drift: Less than 1/800 in frequent earthquake and 1/100 in rare earthquake. Maximum horizontal displacement and drift of the vertical elements should not be more than 1.6 times the average value of the floor when the maximum story drift is no more than 40% of the limiting value under the frequent earthquake.

8.1.7.4 Damping overview 8.1.7.4.1 Damping strategy The half experience and half trial calculation method was used to determine and optimize the damper position. This is based on finding out the weak story with the largest displacement under the earthquake action by a time history analysis for the original structure without dampers. Then dampers are placed where the structural movement is larger and they should also be symmetrically arranged to avoid inducing torsion in the symmetric structure (Figs. 8.67 and 8.68).

8.1.7.4.2 Damping type Sixty two standard viscous dampers were installed, designed, and produced complying with Taylor Devices. The part number of the dampers are 67DP-18900-01, 67DP-18901-01, and 67DP-18902-0. The damper parameters are given in Table 8.18.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.67 Wuhan Poly Cultural Plaza: detail picture of the damping system used.

FIGURE 8.68 Wuhan Poly Cultural Plaza: damper site installation.

8.1.7.4.3 Structural and damping design Code and guidelines. During the design phase the following codes were considered as the basis of design: Code for seismic design of building, GB50011 (2010) Technical specification for concrete structures of tall building, JGJ3 (2010) Analysis modeling and software. The structural analysis model is created in ETABS (CSI, 2016a) finite element program. In the structural analysis calculations the energy dissipation device adopts a nonlinear viscous energy dissipation unit. Design principles. The additional damping ratio that nonlinear viscous dampers bring to the structure is estimated by the method of equivalent contrast under the seismic time history condition. Design phase considerations. The original structure, without damper, was analysed by time history method to find out the weak layer with large displacement under the DE. Then dampers were located where the structural movement was large. Additionally, taking into account the symmetry of the structure, the dampers were arranged symmetrically.

621

Table 8.18 Wuhan Poly Cultural Plaza: Damper Parameter List Type

Damping Coefficient (kN(s/m)0.3)

Speed Exponent

Maximum Damper Force (kN)

Minimum Stroke (mm)

Quantity

67DP-18900-01 67DP-18901-01 67DP-18902-01

2000 2000 2000

0.3 0.3 0.3

1200 1200 1000

6 100 mm 6 75 mm 6 75 mm

6 20 36

Placement Position

Installation Method

8, interlayer 8 37, 39, 41, 43, 44 8, interlayer 8, 22, 24, 31, 33

Chevron Diagonal Diagonal

8.1 Distributed Damping Systems Case Studies

8.1.7.4.4 Architectural integration strategy The overall layout of the building is L-type. The main building is connected with the auxiliary one by a connecting part, and bearings connected them closely and form one structural unit together. The main and the auxiliary building structure adopt round steel tube concrete columns, steel beams, and RC core wall. The buildings are connected between floors 16 and 20 with steel trusses.

8.1.7.4.5 MEP integration strategy There is no interaction between the damping system and the MEP.

8.1.7.4.6 Elevator and other mechanical devices integration strategy Beams below floor 8 adopt H steel section in order to increase the building height and easiness to cross the pipeline equipment. Instead, beams in standard stories use light steel truss beam. Because of the use of the equipment line, the height of the truss beams is larger than H-type steel beam, and thus it can reduce the amount of steel.

8.1.7.4.7 Experimental tests Several experimental tests were conducted for the damper, such as: • •

• •

•

Shape test. Damper and its shape size and appearance are in good shape, no oil leakage, paint spalling, and shell damage. Slow test. Dampers move back and forth slowly at least three cycles, and the relationship between damping force and displacement was recorded. In the case the damper does not leak oil, and the resistance is not beyond the design damping force of 10%, and the basic performance of dampers meets the requirements. Power cycle test. The output force of the damper in the process of a cycle is within the range of 15% of the output force curve of the damper. Pressure test. The hydro-cylinder and piping of the dampers are constantly pressed for 1 hour under a force 1.5 times the safety factor of the design damping. The tested dampers should have no leakage, cracks, or defects. Total stroke test. The total stroke of the dampers meets the requirements of 6 100 mm design value.

8.1.7.4.8 Monitoring system No monitoring system was introduced in the building.

8.1.7.4.9 Maintenance strategy Taylor Devices provided a warranty for 35 years, and a maintenance-free design for the building life, except man-made destruction, being flooded or large earthquake. The damper systems are coated with a fire-retardant paint.

623

624

CHAPTER 8 Case studies of tall buildings

8.1.7.4.10 Development process Wind tunnel test is taken by Structural Wind Engineering Institute of Wuhan University commissioned by the design company (Zhongnan Architectural Design Institute Co., Ltd).

8.1.7.5 Comparison of different design strategies 8.1.7.5.1 Structural options considered This project is a super high-rise building project and belongs to the category of connected structure whose towers are significantly different with a span that is more than 24 m. The National Expert Committee on Seismic Fortification Examination of Overrun High Rise Building Engineering recommended that special review of seismic performance design of this project must be done.

8.1.7.5.2 Damping solution considered The changes of some seismic response index, before and after dampers, are added to the structure as follows: •

•

Base shear and base moment (Table 8.19): for low-level earthquake the energy absorption rate in both X and Y directions varies between 5.37% and 21.95%. For rare earthquake, the ratios range between 0.84% and 7.23%. Vertex displacement: for low-level earthquake the energy absorption rate in both X and Y directions varies between 11.55% and 27.00%. For rare earthquake, the ratios range between 2.59% and 6.41%. Table 8.19 Wuhan Poly Cultural Plaza: Comparison of Base Shear and Base Moment of the Structure Frequent Earthquake

X direction Y direction

Shear (kN) Moment (kNm) Shear (kN) Moment (kNm)

Maximum Maximum Maximum Maximum

No Damper

Set Damper

Shock Absorption Ratio (%)

9167.93 215797.86 6160.84 332694.04

8473.41 204210.60 5524.94 259675.15

7.58 5.37 10.32 21.95

Rare Earthquake

X direction Y direction

Shear (kN) Moment (kNm) Shear (kN) Moment (kNm)

Maximum Maximum Maximum Maximum

No Damper

Set Damper

Shock Absorption Ratio (%)

97800.62 1806026.29 65717.81 3550365.53

96981.72 1740643.86 63903.68 3293570.95

0.84 3.62 2.76 7.23

8.1 Distributed Damping Systems Case Studies

The acceleration peak value of the floor response spectrum is effectively reduced by dampers, thus reducing the seismic response of nonstructural members.

8.1.7.5.3 Cost
benefit analysis Conventional structures are usually designed so that the system meets all strength requirements. Moreover, to meet stiffness requirements additional structure is added, which typically means added cost for steel or concrete. As buildings became taller and more slender, the required amount of additional material becomes larger and larger to achieve sufficient stiffness, and thus adding additional damping system rather than stiffness became more economical and effective. Practice has proved that FVDs, as a structure protection system, can have a better control function and a significant effect by reasonable connection design. During an earthquake, dampers, under moving status, dissipate the seismic energy that the structure will absorb, increase the damping ratio of the structure, decay the vibration process of structure rapidly, and thus protect the main structure against damaging. Damping cost. The total damper cost including products supply, transportation, and erection is: f2,534,200 (US$400,000).

8.1.7.6 Lesson learned and recommendations 8.1.7.6.1 Difficulties in the design The plane layout and elevation of this project is very irregular (Figs. 8.64 and 8.65) and many parts are beyond the limits of the current national standards. Some of the most important irregularities are: •

• • •

The shape and stiffness of the main and auxiliary buildings on both sides of the connected structure, stories 15
19 and 22
46, have a big difference and the connection body plane position is asymmetric, leading to complex connections. The floor slabs under the podium roof of the structure surrounded by three sides. The floor plan with concave and convex is irregular, and the dimension of the concave part of the plane is more than 30% of the corresponding length of its side. The shear bearing capacity of stories 17 and 15, less than the 80% of its upper stories, is a mutation of bearing capacity.

8.1.7.6.2 Design innovative solutions The innovative combination application of nonlinear viscous dampers and BRBs at the central of the connection achieve both goals of reducing and controlling the torsion of the main structure in the earthquake.

625

626

CHAPTER 8 Case studies of tall buildings

8.1.7.6.3 Possible improvements Instead of the design solution utilized in this project, alternative design options to be considered could have been BRBs, buckling steel walls, and other constraints.

8.1.8 TIANJIN INTERNATIONAL TRADE CENTER, TIANJIN, CHINA 8.1.8.1 Project data The major building data (Fig. 8.69) are summarized as follows: • • • •

Year of completion: 2014 Developer/contractor: CapitaLand China Architectural designer: PT Group Structural designer: PT Group

FIGURE 8.69 TianJin International Trade Center: building overview.

8.1 Distributed Damping Systems Case Studies

•

•

• • • • •

•

Damping supplier: Taylor Devices Inc. (United States) Beijing Qitai Shock Control and Scientific Development Co., Ltd Testing laboratory: Damper test by Taylor Devices Inc. (United States) Wind tunnel test by CABR Technology Co. Ltd (China) Height: 235 m from the street level Stories: 60 above grade 1 3 below grade Gross area of the tower: Rectangular plan having dimensions L 5 24 m 3 61 m, total floor area 1.464 m2, total surface 77.600 m2 Building function: Office building Structural material Structural steel: Mainly Q345 Dampers: 67DP-17150-01-1 Floor plan area: Quasi-circular plan having a diameter of 45 m

8.1.8.2 Introduction/history Tianjin International Trade Centre is a skyscraper with 60 floors and a total floor area of 190,350 m2, which is located at 112 Munan Road in Tianjin, China. Tianjin International Trade Centre has a roof height of 235 m. Construction of the project commenced in 1998 but was halted in July 2000. Its construction recommenced on November 25, 2010 and it has topped out on July 11, 2012. It was completed in 2014.

8.1.8.3 Structural system This building is a retrofitted steel structure. All the structural elements including columns, beams, and braces are all steel. The columns have main styles of I-shape, H-shape, and box-shape. I-shapes are mostly used for beams and braces.

8.1.8.3.1 Building fundamental periods The building fundamental vibration modes are shown in Fig. 8.70. The first mode (T1 5 6.02 seconds) is mainly flexural along the Y direction. The second mode (T2 5 5.74 seconds) is mainly flexural along the X direction. The third one (T3 5 2.97 seconds) is mainly torsional.

8.1.8.3.2 Damping strategy utilized FVD is used to reduce wind-induced acceleration. Dampers are arranged with toggle brace to get amplification effect, which could magnify the damper displacement for as much as 2.5 times, thus magnifying the acceleration reduction by the same amount (Table 8.20; Figs. 8.71
8.73).

627

628

CHAPTER 8 Case studies of tall buildings

FIGURE 8.70 TianJin International Trade Center: fundamental frequencies and mode shapes.

Table 8.20 TianJin International Trade Center: Structure Design Parameters Design service life Design reference period Classification for earthquake resistance of buildings Safety classes of building structures Seismic fortification intensity Building site categories Design the basic earthquake acceleration Design of earthquake group Design characteristic period Maximum value of horizontal earthquake influence coefficient Building height

50 years 50 years C-class Second level 7.5 Tianjin III 0.15 g Second 0.63 seconds Frequent earthquake: 0.12; moderate earthquake: 0.34; rare earthquake: 0.72 Over A-class

8.1.8.3.3 Additional damping provided by the damping system The additional damping ratio provided by the toggle brace dampers is estimated to be about 1% by comparing the structural response (in this case story acceleration is used), under the condition of 10-year period of wind case.

8.1.8.3.4 Building cost versus damping cost The damping cost is f1,758,000 (US$280,000), which is far less than the cost of the whole structure.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.71 TianJin International Trade Center: floor framing plans.

FIGURE 8.72 TianJin International Trade Center: section.

629

630

CHAPTER 8 Case studies of tall buildings

FIGURE 8.73 TianJin International Trade Center: elevation picture during construction.

8.1.8.3.5 Building code Several buildings codes were utilized and some others were considered as reference documents. Among all, the most important are: • •

Technical specification for steel structure of tall buildings, JGJ 99 (1998) Technical specification for building with energy dissipation devices, JGJ297 (2013)

8.1 Distributed Damping Systems Case Studies

• •

Dampers for vibration energy dissipation of buildings, JG/T 209 (2007) Code for seismic design of buildings, GB50011 (2010)

8.1.8.3.6 Peer-reviewed project CABR China has done the wind tunnel test for Tianjin International Trade Centre. The consultant of CapitaLand China has made many suggestions on our work.

8.1.8.4 Damping overview 8.1.8.4.1 Damping strategy According to the time history data from the wind tunnel test for the Xiaobailou area of Tianjin, an ETABS (CSI, 2016) model was used to calculate and analyze the comfort level of the structure with the toggle-braced FVDs located in the refuge stories. Dampers parameters and installation locations were optimized to meet the safety and comfort requirements for the structure (Fig. 8.74).

8.1.8.4.2 Damping type The main characteristics of the dampers utilized in this project are as follows: • • • • •

Unit type: double acting FVD Operating ambient temperature range: 0
49 C with minimal change in performance characteristics Damper stroke: 6 100 mm nominally with identical characteristics in each direction of motion Damping force: 979 kN Nominal output function: F 5 CVα (C 5 1200 kN(s/m)α, α 5 0.65)

FIGURE 8.74 TianJin International Trade Center: damper installation.

631

632

CHAPTER 8 Case studies of tall buildings

Table 8.21 TianJin International Trade Center: Damping Effect at the Top Under Wind Load Load Case

Acceleration (m/s2)

Wind

Before damping 0.2237

After damping 0.1992

Decrease rate 11%

Table 8.22 TianJin International Trade Center: Analysis Result of Frequent Earthquake Direction X Story drift Base shear (kN) Base reaction (kN) Base bending moment (kNm)

Without Dampers

With Dampers

Decrease Rate

1/377 25304.89 42244.26 3272.84

1/395 23155.59 41294.53 2985.89

4.4% 8.5% 2.2% 8.8%

8.1.8.4.3 Structural and damping design Code and guidelines. According to the requirements in technical specification for steel structure of tall buildings (JGJ 99 (1998)), the top acceleration should be below 0.2 m/s2 to meet the comfort requirements. Analysis modeling and software. ETABS (CSI, 2016) was used to calculate and analyze the structure with toggle-braced FVDs located in the refuge stories. In ETABS (CSI, 2016), the damping device is modeled by a nonlinear link element named Maxwell model, which is cascaded by a damping element and a linear element (Fig. 5.17 (Chapter 5)), and its total deformation is a contribution by both elements (Tables 8.21 and 8.22). Design principles. Referring to the Chinese code, and experience in similar wind control projects with dampers in the world, the damper design principles are as follows: •

• •

The building, equipped with dampers, should meet the comfort level requirements under the wind load with 10-year return period, and dampers cannot be damaged under wind load with 50-year return period. The dampers designed for wind also need to ensure safety in frequent earthquake and DE, i.e., cannot be destroyed and work properly. Due to the toggle brace and continuous wind excitation, dampers often stroke large displacement motion for a long time, and the designer should also ensure the damper power can fit the continuous work bearing the wind load.

Design phase considerations Appraisal and assessment The upper steel structure has the following problems: too many materials typologies (including varieties imported from different

8.1 Distributed Damping Systems Case Studies

countries), corrosion phenomenon (large quantity and extensiveness, to varying degrees), member deformation, and point corrosion. The construction of some parts do not tally with the drawings: protective coating (corrosion and fire protection) invalidation. Before starting the retrofitting (Figs. 8.75
8.77) a specific implementation plan on the basis of recomprehensive testing was necessary. Derusting and corrosion proof Sand throwing approach is selected with high efficiency, controllable cost, and relatively low potential impact to the surroundings (Table 8.23).

FIGURE 8.75 TianJin International Trade Center: tower retrofit scheme.

FIGURE 8.76 TianJin International Trade Center: floor beam strengthening.

633

634

CHAPTER 8 Case studies of tall buildings

FIGURE 8.77 TianJin International Trade Center: columns and trusses strengthening.

Table 8.23 TianJin International Trade Center: Derusting Technique Comparison Derusting Technique Manual cleaning Sand throwing Chemical washing

Performance

Construction Time

Impact to Environment

Low efficiency at connection joints Good

Long

Minor

Short

Good

Moderate

Controllable with protecting barrier Possible hazard chemical generated

Cost Labor cost mainly Moderate High

Strengthening and seismic proof For floor beams (Fig. 8.76): compensate for the loss in dimension and allow for additional floor load carrying capacities. The contrastive analyses of several plans show that when distributing dampers in brace core, the forces of dampers at top or bottom of the building proved to be large. However, in consideration of the usage of the building, it is decided not to distribute the dampers at the bottom of the building. In addition, when distributing the dampers outside the brace core, the energy-dissipating capacities of the dampers are also remarkable. The configuration of toggle-braced dampers (TBDs) is more complicated than diagonal dampers, and the installation mainly includes pin connection and welding. Due to its particularity, the precision of machining and installing is comparatively high, so the dampers should be guaranteed to be successfully installed and to work effectively.

8.1 Distributed Damping Systems Case Studies

8.1.8.4.4 Architectural integration strategy To avoid the influence to the architectural system and the building function, all the dampers were distributed in the refuge floor.

8.1.8.4.5 MEP integration strategy The dampers are self-sufficient and independent.

8.1.8.4.6 Elevator and other mechanical devices integration strategy The damper system has no influence on other devices.

8.1.8.4.7 Experimental tests The dampers were tested in Taylor Devices’ laboratory; the test items are shown as follows: •

•

Proof pressure test. Each unit is to be proof-pressured to 104,000 13500/ 2 0 kPa for 3 minutes minimum. This corresponds to more than 150% of the pressure produced at full force. No visible signs of physical damage, deterioration, deformation, or fluid leakage are allowable. Sinusoidal performance test (F 5 1200 v0.65 kNs/m 6 15%). Each damper is to be horizontally fixed in a hydraulic tester, and shall be tested with a sine curve for three complete and continuous cycles at 100% of the maximum rated velocity (0.755 m/s). No signs of physical damage, yielding, deterioration, permanent deformation, or fluid leakage are allowable.

8.1.8.4.8 Monitoring system No monitoring system was installed.

8.1.8.4.9 Maintenance strategy Taylor viscous dampers are completely maintenance-free throughout their life cycles, and there is no need for inspection on a regular basis. However, if the owner needs to regularly check for seismic protection reasons, the following procedures are suggested. •

• • •

If the damper is not installed on an accessible frame, a platform should be established near the damper location to visually inspect it at any time. The required number of platforms should be the minimum to check all or at least half of the dampers. The interval of check is about 2
5 years, or after a certain scale earthquake, and there should be a written record of the damper’s condition. Through visual observation, monitor the appearance of the damper to determine whether it operates normally. Damper visual inspection. When stroke or damping forces overload, it is easy to observe one of these: whether there is an oil leakage and whether the damper connections have loosen. No additional checks are required.

635

636

CHAPTER 8 Case studies of tall buildings

Table 8.24 TianJin International Trade Center: Comparison of Three Kinds of Solution for Wind Solution

Explanation

Adjust the stiffness of the structure

Adjust the stiffness of the structure by enlargeing the section of the structure element. Although the acceleration can meet the requirement of the code, the weight of the steel will increase 15%
20% at the same time. Distribute FVDs at the proper place of the structure (adopted plan). Distribute TMD at the top of the structure. Disadvantages: (1) Bad for the vertical bearing capacity of the structure elements and pile foundations. (2) The TMD itself is expensive.

Distributing FVD TMD plan

8.1.8.5 Comparison of different design strategies 8.1.8.5.1 Structural options considered To achieve the goal of improving the comfort level, there are three common solutions: (1) adjust the stiffness of the structure itself; (2) distribute dampers on the structure to increase the damping ratio and decrease the acceleration reaction of the structure; and (3) distribute TMD for vibration reduction. The results of specific researches for the three solutions are given in Table 8.24. From the table, it can be seem that adjusting the stiffness of the structure may need at least 15% extra cost, while using TMD at the top also has many disadvantages such as worsening the structure forces, occupying huge space, and costing expensive. Thus, distributing FVD is the best solution.

8.1.8.5.2 Damping solution considered The structure which has a number of braces in the core from the 1st to the 57th floor is a steel frame-core structure. According to the recommendation of FEMA and the experience of Huntington 111 building in Boston, as a trial scheme, distributed dampers were firstly placed throughout the whole building in the core, and then tested on different plans by using dampers first in every story, then in every other story, followed by 2, 4, or 6 dampers in each story.

8.1.8.5.3 Cost
benefit analysis The dampers’ cost of this project is f1,758,000 (US$280,000), far below the solution of increasing stiffness of the building. Damping cost. Dampers: f1,128,000 (US$179,000) Damper braces and connections: f460,000 (US$73,000) The patent of TBD, reviewing drawings, and relevant assistance: f50,000 (US$8000) Construction drawings and installation guide: f120,000 (US$19,000)

8.1 Distributed Damping Systems Case Studies

8.1.8.6 Lesson learned and recommendations 8.1.8.6.1 Difficulties in the design Strictly speaking, there is no best installing position for dampers because when choosing the positions, the goal is to use the least number of dampers to achieve the best damping effect. However, there is a need to consider whether the damper’s price is too high because of the large stroke and power. Overlarge force of a damper can also make connecting braces cost high or even hard to design. Meanwhile, it is important to consider that some distributing plans can conflict with the function of the building. Therefore, achieving the best distributing plan is very difficult, and usually plenty of distribution plans need to be made for comparison.

8.1.8.6.2 Design innovative solutions The building is the first in China to adopt TBD.

8.1.8.6.3 Possible improvements Aiming at the damper quality problems, the aspects of damper tests must be taken into account seriously. Three pretests in the United States were conducted by HITEC to verify and control the damper performance under tiny velocity, the workshop testing of damper, the calculation method of power and its test, inspection of adjustment, and field detection.

8.1.9 454 YONGE, TORONTO, ON, CANADA 8.1.9.1 Project data The major building data (Fig. 8.78) are summarized as follows: •

• • • • • • • • • •

Year of completion: Start of construction: 2015 Expected top-out: 2017 Occupancy: late 2017 Developer: Canderel Residential (Canada) Contractor: Reliance Construction (Canada) Architectural designer: Graziani & Corazza Architects Inc. (Canada) Structural designer: Read Jones Christoffersen Consulting Engineers (Canada) Damping supplier: Nippon Steel and Sumikin Engineering USA (United States) Damping consultant: Kinetica (Canada) Testing laboratory: Nippon Steel and Sumikin Engineering USA (United States) Height: 200 m Stories: 63 above grade, 5 parking below grade Gross construction Area: 6225 m2

637

638

CHAPTER 8 Case studies of tall buildings

FIGURE 8.78 454 Yonge: building overview.

8.1 Distributed Damping Systems Case Studies

• •

• •

Building function: Residential Structural material Concrete classes: Floors (slab): 35 MPa/42 MPa at 28 days/56 days columns: 80, 70, 60, 50 MPa at 56 days walls/coupling beams (lateral): 75, 65, 55, 45 MPa at 56 days Structural steel: 350 W Dampers: ISD111H viscoelastic material and JIS Rebar: 400R Floor plan area: 18.9 m 3 40.7 m 5 770 m2 Structural system RC coupled shear wall lateral load resisting system Typical levels use structural steel coupling beams and viscoelastic coupling dampers (VCDs) in lieu of conventionally RC coupling beams

8.1.9.2 Introduction/history YC Condominiums is a 63-story residential tower in downtown Toronto. The project is located along a tall building corridor in downtown Toronto on a small site of 38 m 3 45 m bounded by roadways on three sides and an existing heritage structure in one corner of the site. Because of the slender building design (11 to 1 slenderness), frequent lateral wind vibration, which can cause occupant discomfort, was a crucial aspect of the design. The developer’s mandate for this project was to maximize sellable space within a prescribed zoning envelope (total height restricted by shadows and floor plan dimensions restricted by setbacks on three sides and a heritage building on the fourth) which ultimately dictated the choice of damping system for the project.

8.1.9.3 Structural system During the evolution of the project, a number of significant design constraints were introduced; some were market-driven, but many were introduced to suit urban planning considerations due to increasing height of projects in the surrounding area. Structurally, the design constraints required that the tower be tall and slender to accommodate the increased setback requirements and maintain desirable suite layouts. A coupled shear wall lateral load resisting system was employed in which only two primary wall lines were utilized in the narrow plan direction (Fig. 8.80). By concentrating lateral stiffness on these two wall lines, which are located toward the NS extent of the floor plan, it results in reduced torsional mass participation (and torsional velocity) and localized relatively large differential displacements between these walls, which is an optimal condition for incorporating VCDs. This configuration had the added benefits of increased flexibility in suite layouts, while minimizing the number of coupling beams crossing over the corridor (less mechanical coordination). The two primary coupled shear wall lines used structural steel coupling beams on typical tower levels above and below a band of levels with VCDs. These walls

639

640

CHAPTER 8 Case studies of tall buildings

FIGURE 8.79 454 Yonge: fundamental frequencies and mode shapes.

were extended longitudinally within the enlarged footprint of the tower podium (at the lower levels) creating outriggers which bridged loads over drive aisles below grade (Fig. 8.81). The central elevator core and spine walls alongside the residential corridor (in the long plan direction) employed conventionally RC beams.

8.1.9.3.1 Building fundamental periods The building fundamental vibration modes are shown in Fig. 8.79: the first mode (T1 5 6.56 seconds) is mainly flexural along the short direction, the second model (T2 5 4.64 seconds) is mainly flexural along the long direction, and the third one (T3 5 3.84 seconds) is mainly torsional.

8.1.9.3.2 Damping strategy utilized A super-efficient damped structure was realized with 84 identical modular VCD panels, which were produced by Nippon Steel and Sumikin Engineering USA and 3M Japan. Since the VCDs are essentially structural elements, RJC and Kinetica worked together on integrated analysis models and the configuration of the system, with the final damper unit designed by Kinetica. The VCDs (nine international patents and six international patents pending) consist of multiple VE damper panels bolted to cast-in-place steel embeds replacing 42 RC beams on 21 levels of the structure. Fig. 8.83A shows a 3D plan view of a typical floor with VCDs, Fig. 8.83B shows the VCD installed between the primary RC shear wall lines, and Fig. 8.83C shows the wall elevation with the VCDs. Structural steel beams were used instead of concrete coupling beams above and

8.1 Distributed Damping Systems Case Studies

FIGURE 8.80 454 Yonge: plan view.

below the VCDs in order to optimize the coupling stiffness distribution up the height of the building such that more deformation was driven into the VCDs. Structural steel coupling beams had the added benefit of allowing sufficient strength to allow for redistribution of coupling forces when VCD stiffness was ignored during bounding and to standardize the construction technique along this entire line.

8.1.9.3.3 Additional damping provided by the damping system Wind tunnel studies carried out by RWDI (Fig. 8.82) indicated that the building would require approximately 0.9% of supplemental damping to the fundamental mode of vibration to increase the level of human comfort for 1- and 10-year wind-induced motions. Two damping systems were considered by the owner for the project: (1) distributed VCDs and a (2) bilevel TSD tank. Considering the two systems, the building developer selected the VCDs for the project after a detailed technical and financial analysis, which is discussed further herein.

641

642

CHAPTER 8 Case studies of tall buildings

FIGURE 8.81 454 Yonge: structural model.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.82 454 Yonge: wind tunnel test.

643

644

CHAPTER 8 Case studies of tall buildings

FIGURE 8.83 454 Yonge: (A) plan view of building showing VCDs, (B) VCD at coupling locations, and (C) structural elevation A
A.

8.1.9.3.4 Building cost versus damping cost Damping cost was under 1% of total building cost. In addition, the developer increased their revenue which offset the cost of the dampers.

8.1.9.3.5 Building code Several buildings codes were utilized. Among all, the most important are: • • •

The National Building Code of Canada (NBCC, 2010) Design of Concrete Structures: CSA-A23.3 (CSA, 2010a) Design of Steel Structures: CSA-S16-12 (CSA, 2010b)

8.1.9.3.6 Peer-reviewed project A rigorous peer review was conducted by Magnusson Klemencic Associates (MKA, United States) engineers on the damper design and analysis of the damped structure.

8.1 Distributed Damping Systems Case Studies

8.1.9.4 Damping overview 8.1.9.4.1 Damping strategy 3M VE dampers were the first damping system to be used in tall building construction in 1969, when over 10,000 small VE dampers were utilized in each of the World Trade Centers to reduce frequent wind motion. Two other notable and early applications included the Columbia SeaFirst building, with 260 VE dampers, and the Two Union Square buildings in Seattle, with 16 large VE dampers, both built in the 1980s. Since then, over 40,000 dampers have been used in numerous configurations in over 250 projects in some of the most demanding wind and earthquake environments in the world. The VE material currently used is a fourth-generation material which has increased durability, stiffness, and damping characteristics compared to the earlier generation material and is self-bonding to the steel compared to the earlier VE materials that required a bonding agent. The majority of the past 250 projects had VE dampers configured in brace or wall panel configurations to capture racking (or shear-type) deformations, which is an efficient configuration for low-rise buildings and taller steel buildings. Most tall buildings built now are constructed using RC coupled walls, RC core walls, and outrigger-type structural configurations. In these configurations the buildings undergo predominantly flexural deformations inducing relative vertical motion and as such panels and braces that are activated by interstory shear deformations are not as efficient to provide damping unless they utilize displacement amplifying devices. In most current RC buildings, the primary lateral load-resisting systems are formed by vertical structural elements that are coupled to one another using coupling beams or outriggers. When wind and earthquake loads excite the building, these RC members are deformed and stressed heavily in vertical shear. VCDs are introduced in lieu of these RC coupling members between the large vertical structural members making them very effective in providing damping to the overall structure. The VCDs utilize multiple layers of VE material sandwiched between and bonded to multiple steel plates. The VE material alternates between layers of steel plates with each consecutive steel layer extending out and connected to the opposite side. In YC Condominiums, two modular VE damper panels were used at each damped coupling location and bolted using a slip-critical connection to structural steel embeds which were anchored into the concrete walls to form a single VCD (Fig. 8.83B). When the building deforms due to translational or in-plane torsional wind or earthquake vibrations (Fig. 8.84A), the walls rotate and undergo axial deformations, causing the VE material to be deformed vertically in the shear (Fig. 8.84B). This shear deformation causes an instantaneous velocity-dependent force which provides supplemental viscous damping to the building, and an instantaneous displacement-dependent elastic restoring force ensures a coupling effect between the interconnected elements during the dynamic response (Fig. 8.84C). When

645

646

CHAPTER 8 Case studies of tall buildings

FIGURE 8.84 454 Yonge: VCDs under lateral loading (A) deformed shapes of building under acrosswind loading, (B) VE material shear deformations, and (C) VE material hysteresis.

properly configured, VCDs provide highly efficient distributed viscous damping for all types of lateral loads, from frequent low-amplitude wind vibrations to large-amplitude wind storms, and from frequent earthquakes to maximum credible level earthquakes (Fig. 8.85).

8.1.9.4.2 Damping type VCDs were utilized on the project, whereby two VE damper panels are attached to steel embeds which replace RC beams.

8.1.9.4.3 Structural and damping design An integrated VCD design was conducted by RJC and Kinetica for the YC Condominiums. A performance-based design approach was adopted for the design of the VCDs and of the entire structure with VCDs. Code and guidelines. Similar to other buildings, the structural design was complied with all the prescriptions of the National Building Code and RC and Steel Codes in Canada. The following codes were utilized: • • •

The National Building Code of Canada (NBCC, 2010) Design of Concrete Structures: CSA-A23.3 (CSA, 2010) Design of Steel Structures: CSA-S16-12 (CSA, 2010)

8.1 Distributed Damping Systems Case Studies

FIGURE 8.85 454 Yonge: photographs of VE damper panels at manufactures’ facility: (A) 26 VE damper panels, (B) a single damper panel, and (C) view of the VE material and steel layers.

Analysis modeling and software. In ETABS (CSI, 2016), damping was assessed by conducting free vibrational analyses and by measuring the successive peaks of lateral deformation in the short direction with the log-decrement technique using an analytical model with representative serviceability effective stiffness (cracking) properties for the RC members. The dampers were modeled as a spring and dashpot in parallel acting in shear. The optimized configuration of the dampers in the building was achieved by the design team, and the dampers provided over 1% added damping in both the first (primary lateral) and third (torsional) modes of vibration, which are calculated using nominal VE damping properties. After the added damping was estimated, it was provided to RWDI who assessed the human comfort under

647

648

CHAPTER 8 Case studies of tall buildings

the 1 in 1-year and 1 in 10-year loading using the dynamic properties, and the damping will be calculated using ETABS (CSI, 2016). A few important considerations in the design of damping systems are their performance and robustness under expected variabilities in RC tall structures, the effects of this variability on their dynamic response, and the performance of the damping system. These variabilities include the effective stiffness and cracking assumptions of RC members, dynamic modulus of elasticity, and strength of concrete, building mass and inherent damping, and also amplitude dependence of the damping system. Distributed damping systems that do not have amplitude dependence variability are less sensitive to variabilities in the modeling assumptions compared to other damping technologies; however, the design team wanted to investigate this influence on the damping and therefore a sensitivity study was conducted. The study showed that although the natural period of vibration could vary significantly (more than 20%) when possible properties of the structure were considered, the added damping was largely unaffected (less than 10% change in added damping) by changes in amplitude of vibration, modulus of elasticity, and strength of concrete, effective stiffness, building mass, and inherent damping over a range of commonly used assumptions, confirming the robustness of the distributed VCD system. The added damping is actually increased by a reduction in the level of RC effective stiffness (or increase in the level of RC cracking). For this reason, it is anticipated that the level of added damping will increase over time, as the likelihood that building is subjected to stronger loading increases, which leads to cracking, softening of the building, and period elongation (Fig. 8.86).

FIGURE 8.86 454 Yonge: VCD construction sequence: (A) locally fabricated steel sections, (B) fabricated steel elements during construction, (C) elevation view of damper locations, and (D) modular VE damper.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.87 454 Yonge: wind tunnel time histories: (A) rigid pressure model, (B) ETABS model deformed shape, (C) story drift profile, and (D) acceleration response.

Since wind tunnel loading time histories were also provided by RWDI, the lateral accelerations and drifts were also checked in ETABS (CSI, 2016) using a dynamic time history analysis. Fig. 8.87 shows one such result, obtained from a direction of loading that caused particularly large accelerations for a 1 in 10-year loading time history. Fig. 8.87A shows the rigid pressure wind tunnel model, Fig. 8.87B shows a figure of the deformed shape of the building under the wind time history, Fig. 8.87C shows the drift profile of the damped and undamped structure at the corner of the building, and Fig. 8.87D shows the X direction acceleration of the topmost occupied story of the building at a radial distance of 14 m from the central axis of the building. Design principles. The VCDs are a robust distributed damping system, and the dampers reduce all vibrations for both wind and earthquake loadings. Because serviceability performance drove the lateral design requirements, only the beneficial effects of the added damping (and damper stiffness) were conservatively used for enhancing the serviceability performance. For this particular building and configuration, this choice (neglecting VCD damping and stiffness for global strength design) was found to have a negligible effect on the cost of the building’s structure.

649

650

CHAPTER 8 Case studies of tall buildings

Table 8.25 454 Yonge: Example Time History Response for 1 in 10-Year Loading

43 Peak lateral X direction acceleration at a 14 m radial distance (mg) Maximum X direction interstory drift (%) Maximum X direction base shear (kN)

Only With Assumed Inherent Damping

VCD Added Damping

Percentage Reduction

27.6

15.7

43

0.246 8,800

0.125 4,800

49 45

The design consisted of 84 VE damper panels (42 VCDs) at the locations as shown in Fig. 8.83 over 21 levels. Each VE damper panel consisted of nine layers of 5-mm-thick VE material layers bonded in between the steel plates. Photographs of the final VE damper panels produced at the manufacturer’s facility are shown in Fig. 8.85. A summary of the response of the building for this particular wind time history loading conditions is provided in Table 8.25. The peak acceleration, base shear, and overturning moment time history result from ETABS (CSI, 2016) and were very similar to the wind tunnel test results used for design provided by RWDI. It should be noted that the presented data for the undamped building is for only one wind loading input signal and is slightly higher than RWDI reported; however, over more wind tunnel time histories and a typical 1-hour storm duration, the results would be expected to converge to the RWDI results. Design phase considerations. The VE material design bound is an important feature of the VCD design and wind tunnel time history loading was provided by RWDI to evaluate the performance of the dampers under the various design loads. Using the long-term wind loads, an advanced dual mechanical
thermal model developed and validated by Kinetica using more than 500 dynamic tests on VE dampers was used to assess the change in VE material temperature due to selfheating. This was utilized to provide the design bounds for the VE material properties at the natural frequency of vibration with a property modification factor applied to the VE damper damping coefficient to evaluate the added damping provided by the VE dampers. As discussed previously, a conservative approach was used for the strength design, whereby both a lower bound analysis with the VCD damping and stiffness ignored and an upper bound analysis with the damper properties were calculated at a low temperature. The lower bound properties were found to govern the strength design of the building’s structural elements while the higher bound properties were found to govern the strength design of the VE damper panel and its connections. Detailed bounding was conducted to capture the effect of concrete cracking and dynamic modulus of elasticity. This required that several sets of dynamic

8.1 Distributed Damping Systems Case Studies

properties, and their associated damping levels, were provided to RWDI for motion assessment and calculation of the corresponding equivalent static loading. Prior to VCD design process, two baseline sets of properties for a conventional RC building (without VCDs) were provided to RWDI for initial motion assessment and to facilitate TSD design/costing was undertaken (prior to final damping system selection).

8.1.9.4.4 Architectural integration strategy The project went through several design phases with the developer and the architect which resulted in an efficient structural and damping configuration. An important consideration was the construction and installation of the VCDs while accounting for local construction practices in Toronto. During construction of a typical floor, the fabricated steel elements (Fig. 8.86A) are cast-in-place with temporary steel restraint channels with a similar bolt pattern as the VE panels (Fig. 8.86B). After the building is enclosed (Fig. 8.86C), the steel templates are unbolted and removed and the VE damper panels (Fig. 8.86D) are lifted into place (Fig. 8.86E) and bolted to the steel embeds using a slip-critical connection.

8.1.9.4.5 MEP integration strategy Since the VCD replaces only the coupling beams, no major MEP integration for the project was necessary. In fact the noninvasiveness of the VCD system represents one of its most important advantages for building owners.

8.1.9.4.6 Elevator and other mechanical devices integration strategy There was no required integration with elevators or other mechanical devices as the coupling beams were replaced by dampers.

8.1.9.4.7 Experimental tests The VE material used in the VCD has been thoroughly tested and validated against a range of manufacturer properties and the VCD has been tested in full scale, both axially and in a shear configuration. VE material properties are dependent on excitation frequency and material temperature, which is important to quantify in the design. The VE material is only slightly dependent on strain and therefore provides added viscous damping over all strain amplitudes from extremely small micrometer levels to a few centimeters of displacement. This is a very important characteristic and an advantage over many other damping systems which have drastically different performances over different deformation levels. Fig. 8.88 shows a full-scale VE damper test that was carried out at the University of Toronto in a uniaxial configuration at a frequency of 0.15 Hz, with increasing deformation amplitude. As can be seen in the test results the backbone dynamic hystereses are essentially the same from 0.003 mm up to 9 mm displacements through the damper. This very large range of effective damping response enhances the robustness of the system when compared to other damping systems.

651

652

CHAPTER 8 Case studies of tall buildings

FIGURE 8.88 454 Yonge: full-scale VCDs tested axially at a frequency of 0.15 Hz for strain amplitudes 6 0.0003 to 10 mm.

The dampers have been tested in full scale in a shear racking configuration which is representative of the type of deformations they undergo in a real structure (Fig. 8.89A). Fig. 8.89B shows a picture of the VE material deformation during the tests. Fig. 8.89C shows the hysteretic force
displacement response of a harmonic test conducted at a frequency of 0.1 Hz. The same formula presented earlier for the VE material response can be applied to the entire effective VCD force
displacement response in shear usage. If the steel connecting elements are rigid enough, the effective VCD response is essentially the VE material response while if the connection stiffness is relatively low, the effective VCD hysteresis is flatter and dissipates less energy. Careful selection of the stiffness of the steel elements to maintain a high efficiency of the VCDs is an important design parameter. In locations with large seismic demands, a ductile force limiting “fuse” mechanism can be introduced in series with the VE panel portion of the damper such that in the event of an extreme earthquake, the “fuse” members can be capacity designed to activate and limit the forces that are transferred to the structural members. The VCD tested in Fig. 8.90A shows RBSs on both sides of the VE material. With this detail it is possible to achieve substantial shear deformations as a combination of VE material deformations and “fuse” element nonlinear

8.1 Distributed Damping Systems Case Studies

FIGURE 8.89 454 Yonge: full-scale VCDs tested in shear: (A) full test setup, (B) shear deformation in damper, (C) VCD hysteresis at frequency 5 0.1 Hz, and (D) dynamic test causing structural “fuse” members to yield.

FIGURE 8.90 454 Yonge: (A) production of VE material and (B) VE damper tests.

653

654

CHAPTER 8 Case studies of tall buildings

deformations, which is termed a VE
plastic response (see Fig. 8.90B for an ultimate loading test). After an extreme earthquake, the dampers can easily be inspected and if deemed necessary, they could be repaired or easily replaced, thus significantly enhancing the resilience of the system to extreme loading conditions.

8.1.9.4.8 Monitoring system VCDs do not require any specific maintenance, monitoring, or tuning to ensure performance. The reason for this is that the ISD111H VE material has been engineered after 30 years of VE material development at 3M as an inert and stable material. The manufacturer provided (1) data for accelerated aging showing negligible property variation for more than 100 years and (2) test data from real VE dampers that were removed from actual buildings after 20 years in use which showed negligible variability in properties. A fatigue analysis showed that the properties of the VE material were minimally affected by an assumed upper bound lifetime loading of the damper in the building.

8.1.9.5 Comparison of different design strategies 8.1.9.5.1 Structural options considered Because of the width of the site and project slenderness, it was necessary to consider a wind dampening system beforehand. In addition, architectural restrictions made it difficult to simply increase the structural member size or use different structural systems. With these restrictions it was a requirement to implement a damping system on the project.

8.1.9.5.2 Damping solution considered Two damping systems were considered by the owner of the project: (1) distributed VCDs and a (2) bilevel TSD tank. Both damping system designs were taken to an advanced design development level using the same performance targets; detailed costing was then undertaken. The building developer selected the VCDs for the project after a detailed financial analysis by considering the two systems. In addition, the owner required a detailed peer review of the two damping systems conducted by MKA engineers to confirm the selection of the VCD system for this project. Architectural restrictions required that the TSD tank is located eccentrically on the floor plate to suit the layout of rooftop amenity space, and although the intent was to integrate fire suppression into the TSD, the mechanical space was extremely constricted and further stacking of mechanical spaces could not be accommodated at the penthouse levels. The aspects that were considered by the building developer to make the decision are summarized in Table 8.26. Although the above comparison provided the main considerations for the building developer for the selection of the damping system, the building developer should undertake a detailed financial analysis to compare the two systems. Both systems would be able to reliably provide levels of damping for frequent

8.1 Distributed Damping Systems Case Studies

Table 8.26 454 Yonge: Decision Factors by Developer Sellable space

Cost

Construction and quality control

Safety, protection, and design forces

Maintenance plan

Tuning requirements

Because the VCDs do not affect the architecture of the building, the building developer was able to recover over 5000 square feet of space at the penthouse and mechanical locations. TSD and VCD cost was comparable; however, VCD was slightly more expensive at the preliminary costing stage. Both systems were well under 1% of the overall building cost. Installation and construction of both systems are of the critical path of construction for the building. VCD quality control and assurance is largely controlled by the manufacturing process by Nippon Steel and a single VE damper panel design is used throughout the building. The modular VE panels are required to be bolted on-site by local steel installation crew with a weld inspector certifying that the connections have been completed as specified according to the Canadian steel code. Construction of water tanks, waterproofing, etc. is relatively similar to other large water tanks. Quality control and assurance is dictated by the local contractors and the maintenance plan is set up by the consultant. VCD is a distributed damping system and therefore the performance of the structure is positively affected for all dynamic loads including small amplitude frequent wind storms, hurricane level wing storms, as well as all earthquake loads. The TSD induces additional lateral forces at the tank level and additional gravity weight to be designed for, and its performance is not considered as beneficial during loading conditions than serviceability wind loading. VCDs do not require inspection or maintenance to ensure performance, because of the stability of the VE material over the life of the building. The damping provided by VCDs is expected to increase over the life of the structure as it undergoes increased cracking. The modular VE damper panels are located in the corridor and can be accessed for inspection if required. For residential construction in Toronto, the condo board handles the maintenance for TSDs to ensure the water level in the tank matches with the correct frequency of vibration. Also the water in the tank is emptied and the interior area of the tanks needs to be cleaned periodically. VCDs do not require any tuning to ensure performance over the life of the structure. Because VCD is a distributed damping system, the added damping is robust over a wide band of frequencies and loading conditions. Even with the properties of the RC members changing from fully uncracked at the time of construction all the way through to significant cracking levels that are expected for ultimate (Continued)

655

656

CHAPTER 8 Case studies of tall buildings

Table 8.26 454 Yonge: Decision Factors by Developer Continued design level wind loading and moderate to severe earthquake loading over the life of the building, the levels of added damping are largely unaffected by these property changes. TSDs require tuning of the water level at commissioning of the building and can accommodate a variance in the period of vibration. Over the life of the building, if the dynamic properties of the building are affected by changes to RC stiffness properties and/or mass, the water level would need to be adjusted to ensure the effectiveness of the system. The ability to change the water height does offer a level of adjustment as long as the building period of vibration is within a predefined range.

wind vibrations. The main item that came into consideration was the increased profit made using the VCDs, primarily achieved by the sellable space that was recovered at the penthouse level.

8.1.9.5.3 Cost
benefit analysis Building developer underwent a detailed financial analysis comparing a TSD and VCDs. They anticipated a large return for electing to use the VCDs as opposed to TSDs because of the significant additional sellable space at the penthouse level, which was a significant additional revenue. Damping cost. Damping cost was less than 1% of building construction cost.

8.1.9.6 Lesson learned and recommendations RJC, Kinetica, and Nippon Steel and Sumikin Engineering worked closely on the performance-based VCD design and benefited from an extensive and collaborative relationship with RWDI. The structural design followed the National Building Code of Canada (NBCC, 2010) and the Canadian steel and concrete design codes with proper bounding assumptions. The VCD was designed according to the Canadian steel and concrete design codes and followed the recommendations of the JIS (2007) damper design manual with slight modifications for tall building applications. A simplified erection and installation procedure was developed to be efficient and economical using local construction practice in Toronto. Unique slab details were developed for use at the VCD locations to provide additional flexibility in the portion of slab directly over the VCD, to limit cracking and to provide additional erection clearance. Modal damping of the building was calculated through free vibration time history analyses in ETABS (CSI, 2016). VE material properties were modeled in ETABS using a Kelvin
Voigt model (spring and dashpot in parallel (Chapter 5)). The VE

8.1 Distributed Damping Systems Case Studies

material damping and stiffness properties were conservatively bounded in order to confirm the design of the expected VE material properties is achieved fully. Sensitivity studies confirmed the performance and robustness of the damping system over a range of commonly assumed modeling assumptions. The assessment of the wind response by RWDI did not deviate from typical tall building wind studies, with the exception that the damping provided by the VCDs was estimated using the ETABS (CSI, 2016) model with VCDs. As with typical tall building wind tunnel studies, RJC provided the dynamic properties of the building which were used to assess the dynamic wind response of the building. Wind tunnel loading time histories from a rigid pressure model were used to assess the performance of the dampers and further sensitivity studies were conducted. As part of a robust quality control and quality assurance program, prototype and production of VE material and VE damper panel tests were used to ensure consistency between property values used for design and the properties of the manufactured dampers. Because of the demonstrated minimal property variability expected for such a damping system, the VCDs do not require any specific maintenance, monitoring, or tuning to ensure performance over the life of the building. In fact as the building cracks over its lifetime the amount of damping provided by the VCDs increases. The building owner undertook a detailed financial analysis to determine which system to use for their project. A primary consideration for the choice of damping system was the additional sellable space (compared to TSD) that was recovered at the penthouse level. A recommendation by the design team is to work closely with all parties involved early on in the design process as is typical for design of high-profile tall buildings. This also includes working closely with wind tunnel engineers (in this case RWDI) with experience in high-performance damping systems and tall buildings.

8.1.9.6.1 Difficulties in the design Design was guided largely by previous buildings built with VE dampers, in which there has been over 250 guidelines set out by JSSI (2013)—the Japanese Society for Seismic Isolation 2013 Damping guidelines. Because this damper was used for wind comfort, it did involve adapting some reasonable modifications of existing seismic design codes and using full-scale damper tests which had been previously conducted and applied to the design.

8.1.9.6.2 Design innovative solutions The innovative design solution revolves around the damper configuration and replaces an RC structural member where the largest stresses are expected in the building with a high-performance damping system.

657

658

CHAPTER 8 Case studies of tall buildings

8.1.9.6.3 Possible improvements There would be no improvements by a redesign in the damper configuration.

8.1.10 181 FREMONT STREET, SAN FRANCISCO, CALIFORNIA, UNITED STATES 8.1.10.1 Project data The major building data (Fig. 8.91) are summarized as follows: • • • • • • • • • • • • •

• •

Year of completion: 2017 Developer/contractor: Jay Paul Company Architectural designer: Heller Manus Architects Structural designer: Arup North America Ltd (United States) Damping supplier: Taylor Devices Testing laboratory: Taylor Devices Height: 244 m Interstory height: Varies Slenderness ratio: 6.85 Stories: 56 above grade 1 5 below grade Gross area of the tower: 68,263 m2 Building function: Mixes use Structural material Structural steel: Main superstructure and floor framing Concrete: Basement and floor slabs; megacolumns filled with concrete up to level 21 Damper type: 32 viscous dampers, installed in group of 4, each with 440 kips capacity and 6 15.2 cm stroke Floor plan area: Varies with a base width of 35.6 m

8.1.10.2 Introduction/history The 181 Fremont Street Tower, located in Downtown San Francisco, is arguably the most resilient tall building on the West Coast of the United States. It is a mixed-use tower, with high-end residential in the top third of the building and class A office space in the area below. In the heart of San Francisco, adjacent to the Transbay transit center, it is a prime real estate in the city.

8.1.10.3 Structural system The lateral load resisting system is a “dual system” comprising a megaframe on the perimeter of the building with a secondary system using a moment frame within the interior of the building. Within the residential portion of the building, the moment frame is replaced with a braced steel core system, utilizing BRBs. The secondary system acts as a backup to the primary and also to transfer the lateral loads up and down to the meganodes.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.91 181 Fremont Street: building overview.

659

660

CHAPTER 8 Case studies of tall buildings

FIGURE 8.92 181 Fremont Street: typical structural floor plan.

The gravity system (Fig. 8.92) comprises steel floor framing supported by the columns, which in turn are supported by the corner columns via transfer girders. Thus all gravity loading is transferred to the perimeter of the building.

8.1.10.3.1 Building fundamental periods The building fundamental vibration modes are approximately 7.2, 6.6, and 3.4 (torsion) seconds (Fig. 8.93).

8.1.10.3.2 Damping strategy utilized The damped megabrace system, a novel system of providing damping within the length of the megabraces, was utilized (Figs. 8.94 and 8.95).

8.1 Distributed Damping Systems Case Studies

FIGURE 8.93 181 Fremont Street: fundamental frequencies and mode shapes.

8.1.10.3.3 Additional damping provided by the damping system The damping is approximately 8% in two directions for the service level wind. For larger wind events and earthquakes, this effectively increases to 10%
20% because of the nonlinearity in the dampers.

8.1.10.3.4 Building cost versus damping cost This information is not available.

8.1.10.3.5 Building code San Francisco Building Code (SFBD, 2013) is the basis of the design. This refers to other design codes, the most relevant of which is ASCE 7-05 (ASCE, 2005). For the testing of the dampers, ASCE 7-10 (ASCE, 2010) Section 18.9 was used as the main reference.

8.1.10.3.6 Peer-reviewed project This project involves a panel of reviewers worked on behalf of the City of San Francisco, which comprises a mixture of academics and practicing engineers with experience in structural, geotechnical, and related disciplines.

661

662

CHAPTER 8 Case studies of tall buildings

FIGURE 8.94 181 Fremont Street: structural elevations.

8.1.10.3.7 Design forces Information is not available.

8.1.10.3.8 Expected performance Wind acceleration for 1-year wind is 9 mg; for 10-year wind it is 16 mg. Seismic design remains elastic under 475-year event, and it meets REDi Gold requirements (REDI, 2013).

8.1.10.4 Damping overview 8.1.10.4.1 Damping strategy Given the relative slenderness and the weight of the structure (steel framed), the building is relatively susceptible to wind-induced vibration. While typical strategies for reducing accelerations have used TMDs, the location of the building, in a seismic region, favored a robust, distributed damping system. The damping system was required to provide reduction in accelerations for service level wind, which was the major concern with high-end residential

8.1 Distributed Damping Systems Case Studies

FIGURE 8.95 181 Fremont Street: 3D structural model.

663

664

CHAPTER 8 Case studies of tall buildings

apartments at the top of a slender building. In addition, the system needed to provide predictable performance during a range of DEs.

8.1.10.4.2 Damping type A novel damping system was devised for this project. The damped megabrace works by placing a group of viscous dampers at the end of a long bracing element which forms the main lateral resisting system (Fig. 8.96). Since the viscous dampers provide no resistance to static loading, a parallel system of bracing was also used. In the bracing with the dampers attached (the secondary brace), the strain is concentrated within the damper. In the bracing with no damper (the primary brace), the strain was distributed along the length of the damper (Fig. 8.97). This system works best with very long braces; otherwise the stroke that the damper can accumulate is too small.

FIGURE 8.96 181 Fremont Street: megabrace damper 3D representation.

FIGURE 8.97 181 Fremont Street: megabrace damper configuration.

8.1 Distributed Damping Systems Case Studies

Table 8.27 181 Fremont Street: Main Characteristics of Damper Property

Value

Branch 1 (0
5 cm/s) Branch 2 (5
30.5 cm/s) K values (Maxwell springs) Length Stroke

315 kN/(cm/s) 15.8 kN/(cm/s) Minimum 998 kN/cm 2.59 m 6 15.2 cm

FIGURE 8.98 181 Fremont Street: damper and brace installation.

The dampers were placed in eight groups of four within the bracing system. The dampers used were nonlinear, although they were idealized as a bilinear behavior. The dampers were also designed with a smaller stroke than the maximum predicted for the MCE (Table 8.27). In this case the dampers were designed to lock up and act as a strut or tie. The idea was to minimize the volume of fluid within the damper, which would increase the flexibility of the system and hence reduce its effectiveness under service level winds (Fig. 8.98).

8.1.10.4.3 Structural and damping design Code and guidelines. ASCE 7-05 (ASCE, 2005) was used as a primary design code. Analysis modeling and software. Oasys GSA was used as a primary analysis software to calculate static loads within the structure. For nonlinear seismic analysis, LS-DYNA was used to perform NLRHA (Fig. 8.100).

665

CHAPTER 8 Case studies of tall buildings

(A)

(B) 9

40

8

35

7

30

Damping ratio (%)

Roof displacement (in)

Peak displacements Exponent fitted line

25 20

y = 94.5e–0.073x

15 ξ = 8.4%

10 5

6 5

Optimum C value

4 3 2 1

0 0

10

20

30 40 Time (s)

50

60

0

70

0

100

200 300 C value (kips.s/in)

400

FIGURE 8.99 181 Fremont Street: (A) first mode roof peak displacements and (B) damper optimum value. 200 Test

150

LS-DYNA

Dashpot, C Spring, k

100

Force (kips)

666

–1.5

50 0 –1.0

–0.5

0.0 –50

0.5

1.0

1.5

–100 –150

Lockup spring (A)

–200 Stroke (in)

(B)

FIGURE 8.100 181 Fremont Street: (A) LS-DYNA damper modeling and (B) hysteretic behavior comparison.

In order to calculate the level of damping to be assumed during the wind response calculations, LS-DYNA was used to calculate dynamic response. In this case, a simple vibration decay calculation was performed by pushing the building to an appropriate displacement and then measuring the free vibration (Fig. 8.99). The modeled dampers were checked against available test data for dampers used on a similar project (Fig. 8.100). Design principles. Information is not available. Design phase considerations. Information is not available.

8.1.10.4.4 Architectural integration strategy The damper units were hidden behind dry walling within the office space. The size of the dampers was such that they fitted within the envelope of the main bracing elements, so it is not possible to distinguish the dampers from the main steel brace from outside or inside the building.

8.1 Distributed Damping Systems Case Studies

8.1.10.4.5 MEP integration strategy There was no interaction with the MEP systems.

8.1.10.4.6 Elevator and other mechanical devices integration strategy There was no interaction with other elements directly, although the dampers were sized such that they could be fitted within the service elevator if they needed to be replaced.

8.1.10.4.7 Experimental tests Experimental tests were divided into two types—prototype and production. The former underwent a larger number of tests than the latter. The testing regime was based upon that specified in ASCE 7-10 (ASCE, 2010). The three types of tests performed for the prototype units were: • • •

Wind—2000 cycles at the first natural period to test the dampers performance during a long wind storm. Seismic—5 cycles at the maximum velocity were performed. Low velocity checks. The dampers were tested at 0.13, 0.25, 1.27, 2.54, and 5.08 cm/s to verify that the force
velocity relationship was within specification. Testing was performed by Taylor Devices at their production facility.

8.1.10.4.8 Monitoring system There is a monitoring system required to enable monitoring of the building in an earthquake.

8.1.10.4.9 Maintenance strategy The design life of the dampers is 50 years. There is no planned maintenance.

8.1.10.5 Comparison of different design strategies 8.1.10.5.1 Structural options considered The original design did not consider any damping, and it was found that the dynamic performance under service level wind was not acceptable. No feasible solution was found that maintained the architectural intent unless it involved damping.

8.1.10.5.2 Damping solution considered Aside from the damped megabrace, both TMD and TSD were considered. While feasible, these were not pursued, in part because of the high value placed upon floor space at the top of the building.

8.1.10.5.3 Cost
benefit analysis Information is not available. Damping cost. Information is not available. Design implications. Information is not available.

667

668

CHAPTER 8 Case studies of tall buildings

8.1.10.6 Lesson learned and recommendations 8.1.10.6.1 Difficulties in the design Since the design concept requires the braces to dampers to concentrate movement in one location, it is necessary to allow a sliding mechanism that allows the braces to slide relative to the floor beams. This required the use of a “cage” with sliding pads between the brace and the beams. The solution found was effective and simple, but required some thought.

8.1.10.6.2 Design innovative solutions Information is not available.

8.1.10.6.3 Possible improvements Information is not available.

8.1.11 ATUSHI BUILDING, XIN JIANG, CHINA 8.1.11.1 Project data The major building data (Fig. 8.101) are summarized as follows: • • • • •

• • • • • •

•

Year of completion: Not yet completed Developer/contractor: Artux Burak Real Estate Development Co., Ltd (China) Architectural designer: Shenzhen HuaBo Architectural Design Co., Ltd (China) Structural designer: Shenzhen HuaBo Architectural Design Co., Ltd (China) Damping supplier Taylor Devices, Inc. (United States) Beijing QiTai Shock Control and Scientific Development Co., Ltd (China) Height: 75 m Interstory height: 3.0 m/4.5 m Stories: 23 above grade and 2 below grade Gross area of the tower: The total gross area of the tower is 32,476 m2 Building function: Commercial and residential Structural material: Steel: Q345 Steel rebar: HPB300, HPB400 Concrete: foundation: C40 beams, slabs, walls: C30, C40, C50 columns: C40, C50 Dampers: 67DP-19412-01 Floor plan area: Rectangular plan having dimensions L 5 35 m 3 35.5 m, equivalent to a floor area 1,235 m2

8.1 Distributed Damping Systems Case Studies

FIGURE 8.101 Atushi Building: building overview.

8.1.11.2 Introduction/history This project is located in the city of Atushi in Xinjiang. It is a newly built 25-story building, with 2 stories underground and 23 stories above ground. The up ground 1
4 stories are for commercial use, 5
23 stories for residential. The project is still under construction.

8.1.11.3 Structural system The building is an RC frame-shear wall structure; RC frame system composed of beams and columns, shear walls arranged around the plane center and frame in the outer part.

669

670

CHAPTER 8 Case studies of tall buildings

8.1.11.3.1 Building fundamental periods The building fundamental vibration modes are shown in Fig. 8.102. The first mode (T1 5 2.01 seconds) is mainly flexural along the long direction. The second model (T2 5 1.83 seconds) is mainly flexural along the short direction. The third one (T3 5 1.36 seconds) is mainly torsional (Table 8.28).

8.1.11.3.2 Damping strategy utilized For this particular project a series of 56 localized external viscous dampers were utilized to reduce the structure response. The dampers are arranged with toggle brace to get amplification effect (Figs. 8.103
8.105).

FIGURE 8.102 Atushi Building: fundamental frequencies and mode shapes.

Table 8.28 Atushi Building: Structure Dynamic Periods Mode

Period

UX

UY

SumUX

SumUY

RZ

SumRZ

1 2 3 4 5 6 7 8 9 10

2.0083 1.8229 1.3546 0.5551 0.4843 0.3987 0.2761 0.2336 0.2204 0.2203

59.4452 0.1171 0.0002 13.7455 0.1017 0.0125 5.1196 0.0458 0.0002 0.0000

0.1137 53.1995 3.828 0.0554 13.3921 2.266 0.0265 4.9062 0.1443 0.0000

59.4452 59.5624 59.5625 73.3081 73.4098 73.4223 78.5419 78.5878 78.5880 78.5880

0.1137 53.3133 57.1413 57.1967 70.5888 72.8548 72.8813 77.7875 77.9318 77.9318

0.1210 4.5370 43.2713 0.0008 1.0644 10.9801 0.0108 0.2451 0.0229 0.1506

0.1210 4.6579 47.9292 47.9300 48.9944 59.9745 59.9853 60.2304 60.2533 60.4040

8.1 Distributed Damping Systems Case Studies

FIGURE 8.103 Atushi Building: structural plan.

FIGURE 8.104 Atushi Building: toggle brace damper in X direction.

8.1.11.3.3 Additional damping provided by the damping system The additional damping ratio provided by the toggle brace dampers is estimated by comparing the structural response. Under the condition of frequent earthquake, the additional damping ratio in two directions is 14%. Under the condition of rare earthquakes, the additional damping ratio in two directions is more than 3%.

8.1.11.3.4 Building cost versus damping cost Damping cost: f2.576 million (US$410,000)

671

672

CHAPTER 8 Case studies of tall buildings

FIGURE 8.105 Atushi Building: toggle brace damper in Y direction.

8.1.11.3.5 Building code Several buildings codes were utilized and some of them were considered as reference documents. Among all, the most important are: • • • • •

Technical specification for building with energy dissipation devices, JGJ297 (2013) Dampers for vibration energy dissipation of buildings, JG/T 209 (2007) Code for seismic design of buildings, GB50011 (2001) Technical specification for steel structure of tall buildings, JGJ99 (1998) Technical specification for concrete structures of tall building JGJ3 (2010)

8.1.11.3.6 Peer-reviewed project None.

8.1.11.4 Damping overview 8.1.11.4.1 Damping strategy Through mode-superposition response-spectrum analysis and time history seismic analysis, it is computed that the larger story drift is concentrated in-between 10

8.1 Distributed Damping Systems Case Studies

and 25 floors. According to the principle of dampers at positions of large displacement, dampers were placed at floors with larger drift but only at every other story. Specifically, four sets of X direction dampers are placed at stories 11th, 13th, 15th, 21th, and 23th while two sets are placed on the 8th story. Additionally, in the Y direction, four sets of dampers are placed at stories 8th, 10th, 12th, 14th, 16th, 18th, 22th, 24th while two sets are placed on the 3th story. This adds up to a total of 56 damper sets.

8.1.11.4.2 Damping type The damper parameters are as follows (Fig. 8.106): • • • •

Design force F 5 750 kN Damping coefficient C 5 1400 kN/(m/s)0.3 Velocity exponent α 5 0.3 Stroke 6 75 mm

8.1.11.4.3 Structural and damping design Code and guidelines. Several buildings codes were utilized and some of them were considered as reference documents. Among all, the most important are: • • •

Technical specification for building with energy dissipation devices, JGJ297 (2013) Dampers for vibration energy dissipation of buildings, JG/T 209 (2007) Code for seismic design of buildings, GB50011 (2001)

FIGURE 8.106 Atushi Building: FVD drawing.

673

674

CHAPTER 8 Case studies of tall buildings

• •

Technical specification for steel structure of tall buildings, JGJ99 (1998) Technical specification for concrete structures of tall building, JGJ3 (2010)

Analysis modeling and software. The structure model is provided by the design institute. PKPM model was imported into ETABS program (CSI, 2016). This program is used to calculate and analyze the structure during the entire process. The dynamic characteristics of the ETABS structure model were consistent with the PKPM one. FVDs are modeled with nonlinear properties, and nonlinear time history analysis was used. Design principles. Damping was added to keep the structure safe under the 8.5 degree earthquake. The structural response was calculated for the nondamped structure with the damped structure. This is referred to as the equivalent contrast method (Fig. 8.107). Table 8.29 shows dampers offer high damping ratio to the structure. However, the purpose of this design is to ensure that the structure is designed for an

FIGURE 8.107 Atushi Building: comparison of story drift and story shear under frequent earthquake.

Table 8.29 Atushi Building: Damping Ratio Result Summary Level

Direction

Item

Without Damper

With Damper

Damping Ratio (%)

Frequent earthquake

X

Maximum story drift Base shear (kN) Maximum story drift Base shear (kN) Maximum story drift Base shear (kN) Maximum story drift Base shear (kN) Maximum story drift Base shear (kN) Maximum story drift Base shear (kN)

1/550 30364 1/559 27714 1/202 82806 1/205 75579 1/119 140780 1/121 128494

1/967 18348 1/1130 18187 1/282 58305 1/342 56912 1/153 105445 1/175 102830

43.1 39.6 50.5 34.4 28.5 29.6 40.1 24.7 22.4 25.1 31.0

Y Moderate earthquake

X Y

Rare earthquake

X Y

8.1 Distributed Damping Systems Case Studies

FIGURE 8.108 Atushi Building: comparison of story drift and story shear under moderate earthquake.

FIGURE 8.109 Atushi Building: comparison of story drift and story shear under rare earthquake.

8.5 degree earthquake, even though the structure without dampers is designed for an 8 degree earthquake. Therefore, it is necessary to compare the structural response between the structure without dampers under an 8 degree earthquake and the structure with added dampers under an 8.5 degree earthquake. The results are shown in Fig. 8.108. In the case of frequent earthquakes, the vast majority of story shear under an 8.5 degree earthquake, for the damped structure, is less than the original structure without dampers under an 8 degree earthquake (individual points of error of less than 8%). The overall security of the structure under an 8.5 degree earthquake is ensured. Therefore, both goals, reducing project costs and increasing the seismic safety are achieved. The results are shown in Figs. 8.109 and 8.110. Design phase considerations. The dampers were to be arranged as symmetrical as possible.

675

676

CHAPTER 8 Case studies of tall buildings

FIGURE 8.110 Atushi Building: contrast of story shear for the structure without and with dampers under 8.0 and 8.5 earthquake magnitudes.

8.1.11.4.4 Architectural integration strategy The building is allocated for residential and commercial uses. In order to not adversely affect the function of buildings, discussions and consultations were held with the building designers, during the design process, to ensure that the damper design would meet the seismic requirements and at the same time the installation and positioning would meet the architectural and functional demands.

8.1.11.4.5 MEP integration strategy There is no interference between the damping system and the MEP.

8.1.11.4.6 Elevator and other mechanical devices integration strategy There is no interaction between the damping system and the other mechanical devices.

8.1.11.4.7 Experimental tests Before the dampers were delivered, a number of tests were carried out. The damper component level production tests were performed at the Taylor Devices Seismic Test Facility. Each damper was tested to full force and velocity. Additionally, tests were performed to verify the maximum internal pressure capability, temperature tests, and visual inspections.

8.1.11.4.8 Monitoring system A monitoring system is not designed and installed. Taylor Devices provided a warranty for 35 years, and a maintenance-free design for the building life, except

8.1 Distributed Damping Systems Case Studies

in case of man-made destruction, floods, or large earthquakes. The damper systems are coated with a fire-retardant paint.

8.1.11.5 Comparison of different design strategies 8.1.11.5.1 Structural options considered Due to the number of stories of the building as well as the geographic location, the frame
shear wall structure is the priority choice. This design was adopted.

8.1.11.5.2 Damping solution considered During the initial design for adding damping, a BRB system had been considered. However, when comparing BRBs to the fluid damper arrangement, in order to achieve the same damping effect, the BRB cost was higher. Therefore, fluid dampers were selected. The study on the damper’s effect on structures is comprehensive. The patented toggle-brace-damper-connection type can effectively enlarge the damping effect of the damper thereby reducing the number of dampers and ultimately lowering the total cost.

8.1.11.5.3 Cost
benefit analysis Damping cost. Fifty-six toggle brace dampers are used on this building for a total cost of f2.576 million (US$410,000) including all the damper production, transportation, customs’ clearance fee, and installation.

8.1.11.6 Lesson learned and recommendations 8.1.11.6.1 Difficulties in the design The structure is located in an 8.5 degree earthquake zone. High-rise buildings in such a highly seismic region should be designed with a larger cross-sectional dimension beams and columns to meet the needs of earthquake-resistance. The larger the amount of material used, the higher the cost of construction will be. Lowering cost requirements of the owner results in difficulties and challenges to the designer.

8.1.11.6.2 Design innovative solutions The building is the first project in China to adopt the toggle brace damper for seismic control. This toggle brace damper is a patented design by Taylor Devices. It uses the mechanical amplification mechanism to get significant reduction in output force of the damper while achieving the same damping effect, thereby reducing engineering costs. It is an ideal solution for these types of construction projects in such high-intensity seismic zones.

8.1.11.6.3 Possible improvements In terms of costs, the same performance could have been achieved by locating the FVDs in different positions. Force and deformation of local structure should have also been analyzed and calculated, to prevent the damage of the local structure under rare earthquakes.

677

678

CHAPTER 8 Case studies of tall buildings

8.1.12 COSTUMS RESIDENTIAL, AUCKLAND, NEW ZEALAND 8.1.12.1 Project data The major building data (Fig. 8.111) are summarized as follows: • •

Year of completion: Expected completion: 2021 Developer: Shundi Group

FIGURE 8.111 Costums Residential: building overview.

8.1 Distributed Damping Systems Case Studies

• • • • • • • • • • • • •

•

•

Contractor: China Construction Architectural designer: Peddle Thorp (New Zealand) Structural designer: Mott MacDonald Damping designer: Sirve (Chile) Damping supplier: TBD Testing laboratory: Windtech Height: 187 m Interstory height: typically 3.325 m Slenderness ratio: 1/11 Stories: 56 above grade and 5 below grade Gross area of the tower: 46,635 m2 Building function: Luxury residential with podium housing, health club, retail, office, and car parking Structural material Steel-framed tower with composite slabs on metal decking Grade 300 and 350 steel to AS/NZS 3678, 3679, and AS 1163 Total steelwork tonnage B9000 tons (190 kg/m2) Floor plan area: Typical total floor plan area B800 m2 Net floor plan area B680 m2 (85% efficiency) Damping system: A distributed damping system consisting of 28 viscous dampers in a symmetrical configuration of 2-story toggle braces hidden in the stair shaft walls A provisional allowance of a 150-ton TMD at the tapered apex

8.1.12.2 Introduction/history The Customs Residential Project involves the redevelopment of a city central island site into a mixed-use, residential-led development comprising three primary elements: • • •

New Zealand’s tallest residential tower at 56 stories (187 m) with a 7-story mixed-use podium, supported on a 5-level (16-m deep) basement The refurbishment, recladding, conversion, and seismic strengthening of an existing 12-story concrete office building into a boutique hotel The refurbishment and conservation of the heritage “Britomart Hotel”—a public house dating from 1876

As a collection, the buildings across this precinct have been designed to complement each other and provide an offering to the city which is mixed in scale, density, porosity, and character. Together they work to achieve a human scale: delivering an inviting and elegant development which aims to enhance the neighborhood and promote further quality residential development in downtown Auckland.

679

680

CHAPTER 8 Case studies of tall buildings

The urban design drivers require some elegant, integrated engineering solutions. Key features are: •

•

• •

•

•

•

•

This prime piece of waterfront real estate provides a link between the rejuvenated Britomart Precinct and the CBD, with ground level laneways providing activation and permeability. The tower massing responds to the urban context with the historic grain of the site reflected in the primary gridlines and a podium with proportions that reflect the adjacent historic streetscape. The gateway site brings residents back to the heart of the city and provides expansive views over the harbor. To achieve the aspiration of clear unobstructed north-facing views from each apartment, the tower incorporates an architecturally expressed external structural steel megaframe, braced in a diamond pattern that provides a striking form. The megaframe doubles as a key component of the tower’s offset stability system, coupled with the “south core” providing vertical circulation and services reticulation. To respect a neighboring green space, the tower adopts an asymmetric elevation with a tapered apex geometry influenced by the site-shading contours. The shape of the tower is further articulated on the north elevation with a mix of apartments all with north-facing winter gardens and an openable, double fac¸ade and an expressed green feature at duplex apartment levels. The engineering solution also needs to address the constraints of a confined, city central island site to facilitate the consequential phases of conservation, refurbishment, demolition, and tower construction above and adjacent to the five-level deep basement on reclaimed land with a tidal influence.

8.1.12.3 Structural system 8.1.12.3.1 Stability The stability system essentially forms a channel section in plan, with stability in the NS direction provided by the braced megaframes on gridlines 1 and 7 supplemented by outrigger frames on gridlines 2 and 6 and stability in the EW direction provided by a braced megaframe on gridline F behind the south core facilities (Fig. 8.112). The nature of the stability frame geometry leads to an offset between the tower center of mass to its center of stiffness. This leads to a natural drift of the tower to the north during construction—a drift that is significant enough that it needs to be compensated for during construction.

8.1.12.3.2 Gravity system The composite floor slab system has an integrated zone for services with distribution through web penetrations in the fabricated steel beams. The beams are

8.1 Distributed Damping Systems Case Studies

FIGURE 8.112 Costums Residential: typical tower superstructure floor plate with plane arrangement of the stability structural system.

fabricated sections of typically 400-mm deep spanning up to 8.4 m with a 140mm thick concrete slab on metal decking.

8.1.12.3.3 Building fundamental periods The building’s fundamental periods are (Fig. 8.113): • • •

Mode 1: Translation in the X direction (EW); period of 5.41 seconds Mode 2: Translation in the Y direction (NS); period of 4.65 seconds Mode 3: Torsion; period of 2.06 seconds

8.1.12.3.4 Damping strategy utilized Delivering the tower height of 187 m with a structural stability width of 17.4 m (a slenderness ratio of 11) requires an innovative distributed damping system—the first of its kind in Australasia. The tower incorporates a distributed damping system using a collection of 28 viscous dampers in a symmetrical configuration of 2-story toggle braces hidden in the stair shaft walls of the south core between levels 10 and 41 (Fig. 8.114). This damping system has the following key benefits: • •

Discreet arrangement hidden in lines of primary structure leading to no loss in floor area (Figs. 8.115
8.117). Cost-effective using widely available technology that can be competitively tendered.

681

682

CHAPTER 8 Case studies of tall buildings

FIGURE 8.113 Costums Residential: fundamental frequencies and mode shapes.

FIGURE 8.114 Costums Residential: structural elevations with distributed damping system on grids 2 and 6.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.115 Costums Residential: structural model of the tower.

• • • •

Delivers good levels of enhanced damping through the toggle brace amplifier configuration, over several modes of the dynamic response. Viscous dampers provide effective additional damping for very small vibrations as well as for large displacements. Distributed and resilient with multiple devices that do not require tuning. Reliable, airtight sealed, and do not require maintenance.

In addition to the distributed damping system the design also includes for the provisional allowance of a 150-ton TMD at the tapered apex of the tower.

683

684

CHAPTER 8 Case studies of tall buildings

FIGURE 8.116 Costums Residential: typical 4-story structural model.

FIGURE 8.117 Costums Residential: architectural visualization of the toggle brace viscous damping system.

The need or otherwise for the incorporation of the TMD will be assessed during construction once the actual building dynamic characteristics are measured and compared to those that have been predicted by the analytical assessment.

8.1.12.3.5 Additional damping provided by the damping system Different damping configurations were studied and the relative predicted additional damping provided is given in Table 8.30.

8.1 Distributed Damping Systems Case Studies

Table 8.30 Costums Residential: Additional Damping Provided by the Damping System Equivalent Modal Damping Ratio (ξ) X Wind Direction (300 degrees) Damper Configuration 1: 28 VDs 2: 36 VDs 3: 36 VDs 1 TMD

Y Wind Direction (30 degrees)

Mode 1 (X)

Mode 2 (Y)

Mode 3 (Rz)

Mode 1 (X)

Mode 2 (Y)

Mode 3 (Rz)

1.6% 1.7% 4.8%
5.9%

2.2% 2.7% 3.8%
4.6%

2.9% 3.0% 2.4%

1.7% 1.9% 4.9%
5.9%

2.3% 2.9% 3.9%
4.9%

3.8% 3.9% 3.2%

VDs, viscous dampers.

Configuration 1 was selected as the preferred configuration from a cost, performance and architectural perspective, noting that should an enhanced performance be desired at a later point in time—the incorporation of a supplemental TMD would deliver this.

8.1.12.3.6 Building cost versus damping cost The building construction cost is estimated at approximately $300 million and distributed damping cost is approximately $500 million. This is less than 0.2% of the total building cost.

8.1.12.3.7 Building code New Zealand Building code: AS/NZS 1170 (AS/NZS, 2011) and NZS 3404 (NZS, 2009).

8.1.12.3.8 Peer-reviewed project Aurecon is the designated peer review team for the project.

8.1.12.3.9 Design forces Three demand levels were considered for the design of the supplemental damping system: • • •

Service wind demand level given by the 1-year return period wind load Ultimate wind demand level given by the 1000-year return period wind load Ultimate seismic demand level given by the 2500-year seismic spectrum and compatible accelerograms

8.1.12.3.10 Expected performance The target performance criteria for occupancy comfort follows recommendations by Melbourne (1998), ISO 10137 (ISO, 2007), and Isyumov (1995). These include maximum values for peak accelerations and rotational velocities for a 1-year return period wind, as indicated in Table 8.31.

685

686

CHAPTER 8 Case studies of tall buildings

Table 8.31 Costums Residential: Maximum Acceleration Criteria Type

Criterion

Maximum Value

1-year return period peak acceleration (minimum objective) 1-year return period peak acceleration (ideal objective) 1-year return period peak rotational velocity

Melbourne (1988)

10.2 mg

ISO 10137 (ISO, 2007) for residential towers Isyumov (1995)

7.7 mg 1.5 mrad/s

8.1.12.4 Damping overview 8.1.12.4.1 Damping strategy The tower superstructure incorporates a novel distributed damping system that is designed to control the building sway and accelerations within reasonable limits, with specific criteria for wind loads and occupancy comfort acceleration limits. A stochastic wind load model has been developed to estimate acceleration performance for different damping configurations. The distributed damping system consists of an arrangement of viscous dampers that have been added to the building structure in discrete locations up to the height of the tower, generating supplemental damping to several modes of the dynamic response of the building (more than just the principal modes). Should further supplemental damping be required to reach desired performance objectives, the design has also considered the addition of a TMD at the tapered apex of the tower. Whether this TMD is required or not will be evaluated during the construction phase by monitoring the actual dynamic behavior of the building. Space has been reserved for the location of a 150-ton nested pendulum near the tapered apex of the tower. The concept design of the TMD has been completed and its supplemental performance has also been assessed. The nested pendulum system has been designed to be tuned to two different frequencies of the building, the main modes in each direction (X and Y). To increase energy dissipation capacity of the distributed damping system, nonlinear viscous dampers were considered. Relative deformations between connected floors are small (0.35
0.6 mm root mean square (RMS)) for occupant comfort (OCC) wind loads. Displacements in this range require special type of viscous dampers and the system might be vulnerable to other efficiency losses. As the deformations on dampers are small, any loss in elastic deformations and gaps might generate an inefficient performance. Toggle mechanisms (Fig. 8.118) can increase damper deformations by about 5 times, making them work at more conventional stroke levels and smaller forces. Both of these effects lead to more cost-effective devices. This evaluation dictated alternative arrangements considering the use of toggle braces connecting two consecutive stories to increase damper deformations and deformation rates. The toggle connection design needs to consider a system to override the structural elements at the mid-story as shown in Fig. 8.123. The kinematics of the

8.1 Distributed Damping Systems Case Studies

FIGURE 8.118 Costums Residential: details for an override arrangement (to bypass the beam at the midstory of the toggle system).

toggle mechanism and the relative displacement amplification, ratio of relative displacements Δy/Δx, as a function of the initial angle θo of the axially rigid braces are shown in Fig. 8.119. The displacement amplification Δy/Δx . 1 is achieved for any brace angle θo , 26.5 and it increases for smaller angles. For θo 5 6 , amplification of the order of 4.75 is achieved. The damping system considers 28 viscous dampers located on two gridlines of the building structure, 14 on GL2 and 14 on GL6. Fig. 8.120 shows the arrangement of the damping system on GL2. If required, a 150-ton TMD system will be located at level 55 of the building, as shown in Fig. 8.121. The following aspects were considered for the preliminary design of the TMD system. The available height for the TMD in level 55 is limited. Thus, the TMD system proposed is based on a “nested TMD” solution, which produces higher pendulum lengths (a double pendulum system). The available space for the TMD on level 55 requires the design of a longshaped mass hanging from the double pendulum system. Thus, it will potentially require an anti-jaw mechanism. The weight considered for the TMD is 150 tons, which improves the reduction of wind-induced vibrations. The preliminary target periods of vibrations of the TMD system, and the required pendulum length for each direction of the building are X direction: Tx 5 4.86 seconds. Pendulum length: 5.87 m. Y direction: Ty 5 4.40 seconds. Pendulum length: 4.80 m. For fine-tuning of the TMD system, the pendulums are designed to be capable of adjustment in each direction (independently).

687

688

CHAPTER 8 Case studies of tall buildings

FIGURE 8.119 Costums Residential: toggle mechanisms kinematics and displacement amplification as a function of toggle brace angles.

Four viscous dampers are considered for optimal damping of the TMD system. These dampers are located in inclined positions with respect to the horizontal plane, producing smaller strokes and higher loads in the devices. The required displacement of the TMD for 1-year return period wind is about 100 mm. The TMD geometry is designed to achieve 500 mm of maximum lateral displacement. Figs. 8.122
8.124 present the preliminary design of the TMD system that achieves all the design considerations listed above.

8.1.12.4.2 Damping type The viscous dampers nominal properties are: • • •

c 5 50 kN(s/m)0.30. α 5 0.30. Minimum stroke 5 150 mm.

FIGURE 8.120 Costums Residential: building elevation with distributed damping system (toggle brace and viscous dampers).

FIGURE 8.121 Costums Residential: provision for a 150-ton TMD at level 55.

690

CHAPTER 8 Case studies of tall buildings

FIGURE 8.122 Costums Residential: transverse view of the proposed TMD system at center position (A) and at maximum displacement (B).

FIGURE 8.123 Costums Residential: longitudinal view of the proposed TMD system at center position (A) and at maximum displacement (B).

8.1 Distributed Damping Systems Case Studies

FIGURE 8.124 Costums Residential: upper floors elevation view with location of the potentially required TMD.

Occupant comfort wind loads (1-year return period). Velocity, stroke, and dissipation power required for the viscous dampers are the following: • • • • •

Damper velocity (min): 2.5 mm/s Damper velocity (max): 9.05 mm/s Damper dynamic stroke (half amplitude-min): 6 2 mm Damper dynamic stroke (half amplitude-max): 6 6.4 mm Damper dissipation power required: 15.0 (Nm)/s

Ultimate limit state wind loads (1000-year return period). Velocity, stroke, and dissipation power required for the viscous dampers are the following: • • •

Damper velocity (max): 60.27 mm/s Damper dynamic stroke (half amplitude-max): 6 51.3 mm Damper dissipation power required: 230.6(Nm)/s

691

692

CHAPTER 8 Case studies of tall buildings

Ultimate earthquake loads. Velocity, stroke, and dissipation power required for the viscous dampers are the following: • • • •

Damper velocity: 777.12 mm/s Damper total dynamic stroke: 1107.6 mm, 166.3 mm (excluding deformations from static loads) Damper maximum force under ultimate limit state (ULS) demands: 53.3 kN Damper dissipation power: 3,053.3 (Nm)/s

8.1.12.4.3 Structural and damping design Code and guidelines. The structural design complies with: • • • • • •

NZ Building Code B1/VM1 AS/NZS 1170 (AS/NZS, 2002) Structural Design Actions (all parts) NZS 3101 (NZS, 2006) Concrete Structures Standard AS/NZS 4671 (AS/NZS, 2001) Steel Reinforcing Materials NZS 3404 (NZS, 1997) Steel Structures Standard AS/NZS 1154 (AS/NZS, 2014) Structural Steel Welding (all parts)

Analysis modeling and software. The process of damper configuration design and damper parameter definition required the development of an analysis methodology and customized software that could estimate stochastic performance of the structure for 1-year return period wind events, given the nonlinear constitutive relation of the viscous dampers. Stochastic analyses of a linear equivalent structural model were performed with nonlinear distributed dampers, using statistical linearization techniques and a stochastic wind load model. A reduced order model of the structural system was developed by exporting from ETABS structural model (CSI, 2016), and 200 mode shape vectors computed using load-dependent Ritz vectors associated with forces on elastic bracings connecting each damper to the main structure for each damper configuration analyzed. The methodology is conceptually depicted in Fig. 8.125. A stochastic model for “along the wind,” “across the wind,” and torsional moments acting at each level of the structure was developed using theoretical formulations available in the literature. The main characteristics of the wind load model developed for this project are the following: • • • • •

Random characterization of wind turbulence. Includes spatial coherence of wind forces through spatial coherence functions of wind turbulence. Represents stationary demand along the wind, across the wind, and torsional effects of wind pressures. The model uses wind information based on wind tunnel tests (provided by Windtech). The stochastic load model provides also artificial wind load signals for validation with a nonlinear response using conventional structural software (ETABS (CSI, 2016)).

8.1 Distributed Damping Systems Case Studies

FIGURE 8.125 Costums Residential: conceptual representation of the methodology for damper design.

The stochastic load model for X and Y wind directions assumes statistical independence of along-the-wind, across-the-wind forces, and torsional moments applied at each level of the structure. The model is defined in the frequency domain by the corresponding cross power spectral density (PSD) matrices for the along-the-wind force vector, across-the-wind force vector, and the torsional moment vector, as functions of frequency (in Hz). Base moments RMS and PSDs were used to fit parameters of the stochastic force model. The stochastic wind force model was checked and adjusted computing the PSDs of the base moments of the model and comparing them with experimental RMS and PSDs of base moments in X, Y, and Z directions to enforce the same RMS base moments as those reported by the wind tunnel testing in each wind direction considered. The result of this adjustment for the X and Y wind directions is presented in Fig. 8.126, where the thin blue line is the experimental result PSD in log10 scale, and the thick green line is the PSD of the theoretical wind model. The general process for developing the probabilistic analysis and obtaining equivalent modal damping and performance results is shown in Fig. 8.127. Because equivalent damping parameters depend on RMS deformation rates of the dampers and deformation rates of the dampers depend on the equivalent damping matrix where the equivalent damping parameters are assembled, the solution of the stationary response of the nonlinear model using statistical linearization requires an iterative procedure. Once convergence is achieved, mean square response of different quantities of interest can be computed, such as damper deformation rates, floor accelerations at different locations, and floor angular velocities.

693

694

CHAPTER 8 Case studies of tall buildings

FIGURE 8.126 Costums Residential: conceptual representation of the methodology for damper design.

FIGURE 8.127 Costums Residential: summarized flowchart for probabilistic wind analyses in the building Y direction.

8.1 Distributed Damping Systems Case Studies

For validation purposes, a complete structural model (ETABS (CSI, 2016)) was analyzed considering nonlinear constitutive relationships of proposed VDDs, and a set of time history wind load signals compatible with the stochastic wind model used in reduced order analyses. The purpose of generating random samples of wind forces and moments compatible with the developed wind model was to validate peak damper forces and deformations, peak total acceleration, and peak angular velocity reductions at level 52 using numerical simulation in the nonlinear models. Fig. 8.128 presents the results obtained for the comparison between probabilistic and time history analyses. In all cases the differences between both models are less than 10%, which is considered reasonable for results coming from totally different analysis approaches. Design principles. The main goal of the distributed damping system is to achieve the 10.2 mg maximum acceleration limit, considering the 1-year return period wind load (OCC). A cost-performance optimization was carried out comparing different dissipation configurations. Performance was measured in terms of modal damping increase and acceleration performance. A parametric analysis was performed modifying the damping constant of the toggle dampers in the range where the optimum value that maximizes equivalent modal damping is located (starting from 20 kN(s/m)α to a maximum of 80 kN (s/m)α) for different damper arrangements. An example of the results of equivalent damping and mean peak damper forces obtained for one of the damping configurations is shown in Fig. 8.129.

FIGURE 8.128 Costums Residential: comparison of probabilistic and nonlinear time history analysis.

695

CHAPTER 8 Case studies of tall buildings

Y-wind (30 degrees) 0.06

0.06 Model damping ratios ξ1 (b) : ξ2 (g): ξ3 (r)

Mode 1 (X) Mode 2 (Y) Mode 3 (Rz)

0.05

0.04

ξ3 (Rz) = 3.0% 0.03

ξ2 (Y) = 2.7%

0.02

ξ1 (X) = 1.7%

0.01

0

Mode 1 (X) Mode 2 (Y) Mode 3 (Rz)

0.05

ξ3 (Rz) = 3.9%

0.04

ξ2 (Y) = 2.9%

0.03

0.02

ξ1 (X) = 1.9% 0.01

0 2

3

4

5

6

7

cα [N(s/m)0.3]

40

2

x 104

35

35

30

30

25 Diagonal dampers

20 15 Toggle dampers

10

3

4

5

5

6

7

6

7

cα [N(s/m)0.3]

40

Peak damper forces [kN]

Model damping ratios ξ1 (b) : ξ2 (g): ξ3 (r)

X-wind (300 degrees)

Peak damper forces [kN]

696

25

x 104

Diagonal dampers

20 15

Toggle dampers

10 5 0

0 2

3

4

5

cα

[N(s/m)0.3]

6

7

x 104

2

3

4

5

cα [N(s/m)0.3]

x 104

FIGURE 8.129 Costums Residential: parametric evaluation of equivalent modal damping and damper forces for different values of damping constant in toggle dampers.

Design phase considerations. The design of the distributed damping system occurred during the detailed design phase of the project, with Sirve and Windtech both acting as specialist design sub-consultants to Mott MacDonald. A close interaction was required between the Mott MacDonald structural engineers, the specialist damping design engineers from Sirve, and the wind engineering consultants at Windtech. The preliminary and detailed design was complete during a period of 6 months.

8.1.12.4.4 Architectural integration strategy The toggle system is proposed between the building primary columns over a double-story height. The system is incorporated in a symmetrical arrangement on gridlines 2 and 6, between level 10 and level 41 in the tower (Fig. 8.125). The benefit of locating the toggle brace dampers in this configuration is that there is minimal effect on the tower architecture and they are easily coordinated with the building services.

8.1 Distributed Damping Systems Case Studies

8.1.12.4.5 MEP integration strategy With the location of the distributed damping system hidden within the stair shaft walls, there is a very limited effect on the building services components. Apart from leading to an isolated zone for small power conduit, there are very few other elements that are affected.

8.1.12.4.6 Elevator and other mechanical devices integration strategy The incorporation of the distributed damping system did not lead to any coordination issues with the elevators or any other mechanical equipment.

8.1.12.4.7 Experimental tests Production tests. Production testing (quality control testing) will be done to every device to be installed on the building (100% testing). Each device will be tested separately through a sequence of 5 sinusoidal cycles at the frequency given in Table 8.32. Prototype tests. Two devices will be subject to prototype testing, considering the requirements of Table 8.33.

8.1.12.4.8 Monitoring system One-year return period wind vibrations. For an adequate and robust measurement of accelerations (translational and torsional) the monitoring system requires at least three sensors to be installed: •

One sensor on the level 52 floor slab in an accessible location within the landlord space Two sensors on the level 41 floor slab to be located in accessible locations

•

Stronger wind conditions and seismic events. Two additional accelerometers shall be installed for vibration measurement under seismic and stronger wind events
one in the basement of the building and one in level 52. Table 8.32 Costums Residential: Production Tests Parameters Type

Number of Cycles

Frequency (Hz)

Amplitude (mm)

VD1

5

0.22

64

Table 8.33 Costums Residential: Parameters for Prototype Tests Phase

Type

Number of Cycles

Frequency (Hz)

Amplitude (mm)

A B C

VD1 VD1 VD1

5 2000 2000

0.82 0.19 0.24

6 150 6 20 62

697

698

CHAPTER 8 Case studies of tall buildings

The monitoring process will be performed in three main stages: •

•

•

Stage 1: Measuring the inherent damping ratio and natural frequencies of the structure (without any vibration reduction system). During the construction phase, two biaxial accelerometers (X and Y direction) will be installed at level 41 and one on level 52. At this stage, the dynamic parameters of the building without any vibration reduction system installed will be calculated using the method proposed by Jann et al. (2004). This method can give an accurate prediction for the response of tall buildings using ambient wind vibration data providing the information about the in situ dynamic properties of the building, including the natural frequencies and damping ratios. Since three accelerometers will be used, the method will allow estimating the dynamic parameters from three different sources, and by comparing these measurements and the building finite element model response, the dynamic properties of the three first modes of the building that generate the movement can be estimated. Once the dynamic parameters are obtained, the designer will use these values to recalculate the performance of the building under the serviceability limit state (SLS) wind vibration conditions to evaluate if the desired performance is achieved with or without a TMD. From these results a final decision regarding the use of a TMD can be taken at this stage. Stage 2: Measuring the damping ratio and natural frequencies of the complete system including the vibration reduction system. Once the vibration reduction system is installed, the dynamic parameters will be calculated using the same method described in Stage 1 to validate the results obtained from the finite element model computed using the parameters measured in Stage 1. Stage 3: Permanently monitoring the building response. Once the construction of the building is completed, the same three biaxial accelerometers used in Stage 1 and Stage 2 will remain installed with the purpose of measuring the performance of the building permanently. These sensors will be capable of measuring small vibrations originated by the 1-year return period wind conditions. Additionally, two triaxial accelerometers for stronger movements will be installed in the building: one at the foundations level and the other at the top of the building (level 52). These sensors are capable of measuring strong motions related to both a 1000-year return period wind vibration and strong seismic movements. The location of these sensors has to be carefully selected. They have to be protected from the environment and the construction process risks and should be easily accessible. These sensors will measure acceleration constantly, and the system will record in a local memory an event when a predefined acceleration level is exceeded (triggering). The gathered information has to be remotely accessible via Internet in order to analyze the data and obtain the dynamic parameters of the system whenever needed.

8.1 Distributed Damping Systems Case Studies

8.1.12.4.9 Maintenance strategy The viscous damper devices are “airtight” and they are not expected to require any maintenance for the design life of the building. The devices are to be supplied with a warranty of 15 years. A monitoring system is proposed for measuring the performance of the damping system and periodic inspections of this monitoring system will be conducted, typically on an annual basis. The distributed damping system will be fire-protected like any other element of the steel structure to a fire-resisting rate of 60 minutes.

8.1.12.5 Comparison of different design strategies 8.1.12.5.1 Structural options considered The design development of the structure considered numerous options at the conception, preliminary, development, and detailed design stages for the tower superstructure. These included: • • • • • • • •

Braced RC cores to the south elevation External steel-braced megaframe Dual internal steel outriggers Shoulder dampers (TMD or tuned liquid damper, TLD) Apex damper (TMD or TLD) South core external braced frame Supplemental sway frame Distributed VE damper system 1 double column

Each design stage also included a development in the rationalization of the integrated design of the primary structure which is architecturally expressed.

8.1.12.5.2 Damping solution considered Three principal options for the damping solutions were considered for the tower: • • •

TMDs or TLDs at the shoulders and at the apex of the tower A distributed system of VE dampers positioned between a double-column arrangement on gridlines 2 and 6 A distributed system of viscous dampers in a toggle-braced configuration over 2 stories on gridlines 2 and 6

8.1.12.5.3 Cost
benefit analysis A high-level cost
benefit analysis was completed on the three primary damping systems proposed, identifying the one clear favorite which was adopted. Shoulder 1 apex dampers (tuned mass damper or tuned liquid damper). The principal shortcoming of this option was that a collection of three TMDs or TLDs

699

700

CHAPTER 8 Case studies of tall buildings

were proposed in order to provide the supplemental damping for the three primary modes of the tower (both translation and torsion). The asymmetrical tapered apex of the tower also led to inefficiency in the placement and design of these discrete supplemental dampers with each being located on a different floor in the tower. The space that was initially considered as being required for these dampers would have led to a net reduction in internal area of approximately 250 m2 and compromised the sale of two units in the tower. This is a loss in revenue of approximately US$2.9 million. For this specific reason, it was quickly discounted. Viscoelastic distributed damping system. Advice on initial design was provided which led to the adoption of this solution during detailed design. The solution has the benefit that it did not occupy net salable area in the tower, and could be hidden in lines of structure on gridlines 2 and 6, it can reliably deliver the supplemental damping required, it does not require tuning, it does not influence construction speed, and it does not require maintenance. The principal detractors though were that once designed and specified there is limited commercial competition in the manufacture and fabrication of the devices, and the fact that a double-column configuration was required in the building between which the devices would be located. The budget design fee for the VE damping system was US$130 million, and the budget cost for the fabrication, supply, and monitoring of the devices was US $1.3 million. Viscous toggle brace distributed damping system. Following a cost
benefit analysis during the detailed design phase, the tower superstructure was modified to reflect the adoption of this solution. This needs to consider some of the abortive costs associated with changing the structural configuration away from the double-column arrangement which has already been coordinated. The advantages of the distributed viscous damping system were much the same as that for the VE distributed system, and from a technical perspective there was a greater confidence in their performance since they are velocity dependent and perform well for both small and large deformations of the tower. In addition to this there is a commercial competition in the manufacture and fabrication of the devices following design and specification. This is a significant benefit as the total budget estimate for the devices themselves is in the order of US$400 million, even though the design fee was much the same as the VE system. Damping cost. The total design costs for distributed damping and specialist wind consultancy came to approximately US$150 million. The budget for the manufacture, fabrication, and supply of the devices is US $400 million (excluding the steel elements of toggle brace configuration). The budget for the testing, commissioning, and installation of the monitoring system is US$100 million.

8.1 Distributed Damping Systems Case Studies

8.1.12.6 Lesson learned and recommendations 8.1.12.6.1 Difficulties in the design Difficulties encountered during the design were principally due to the late selection of the preferred supplemental damping system. This late adoption of the distributed viscous toggle brace system required the following modifications to the primary structure: • • •

Removal of the double-column arrangement which had previously been coordinated for the VE damping system Reconsideration of the story stability system that connects back to the megaframe which is on a double-story module Development of the double-story toggle brace “override” device that moves independently from the primary structural frame

8.1.12.6.2 Design innovative solutions The innovative solutions adopted with this design include: • •

•

A 2-story toggle brace system to provide an amplification to the performance of the viscous dampers Designing the location of the distributed damping system to be on two primary structural gridlines which is easily accessible (from within the stair shafts) and also contributes to the architecture of the building The adoption of a probabilistic analytical approach to determine the optimum damper configuration, the optimum damper properties, and to deliver a robust design

8.1.12.6.3 Possible improvements The principal lessons learnt from the design and implementation of the damping system are: • • •

Consider the cost
benefit analysis of different supplemental systems earlier in the design process. Consider the whole cost of damping systems and seek where possible commercial competition in the design, manufacture, and supply of devices. Understand how the incorporation of the damping system influences the gravity performance of the structure and load paths in the ULS.

8.1.13 CONNOR TOWER, MANILA, PHILIPPINES 8.1.13.1 Project data The major building data (Fig. 8.130) are summarized as follows: • • •

Year of completion: Expected completion: 2020 Developer: Ortigas and Company (Philippines) Architectural designer: GF & Partners Architects (Philippines)

701

702

CHAPTER 8 Case studies of tall buildings

FIGURE 8.130 Connor Tower: building overview.

8.1 Distributed Damping Systems Case Studies

• •

• • • • • • • • •

• •

•

Structural designer: Sy^2 1 Associates, Inc.—Structural Engineers (Philippines) Damping supplier: Nippon Steel and Sumikin Engineering Co. (Japan) PNS Advanced Steel Technology (Philippines) Damping consultant: Kinetica (Canada) Testing laboratory: Nippon Steel and Sumikin Engineering Co. (Japan) Height: 174.5 m from the base Interstory height: 2.9 m Slenderness ratio: 6.5:1 (NS direction) and 3.6:1 in (EW direction) Stories: 57 above grade Gross area of the tower: 87,587 m2 Building function: Residential with retail stores in the podium Structural material: Concrete classes: • Floors (slabs): Varies from 35 to 41 MPa • Columns: Varies from 41 to 69 MPa • Walls/coupling beams: Varies from 41 to 69 MPa Structural steel: ASTM A992 Floor plan area: 48 m 3 27 m 5 1296 m2 Structural system RC coupled shear wall system in one direction (EW direction) and RC outrigger system in the other direction (NS direction). VCDs are used in middle outrigger system. Damping system: VCD

8.1.13.2 Introduction/history Connor tower is a 57-story residential building located in San Juan, Metro Manila, Philippines. The typical floor plan of the building is 48 m 3 27 m. The slenderness ratio along the flexible axis of the building is 6.5 to 1. The building is located in an area with strong winds including typhoons as well as strong seismicity with the West Valley Fault capable of producing an earthquake with estimated magnitude of 6
7 and greater, running less than 3 km from the site. This required a robust structural system for the building not only to reduce wind occupant comfort, drifts, and design wind loads, but also to increase its resilience against earthquakes. Based on the occupancy, the building falls in seismic risk category II. Since the building falls outside the code-prescribed limits for seismic design, a performance-based seismic design (PBSD) approach was used (PEER-TBI, 2017; LATBSDC, 2017). Due to high flexibility in the short direction, an RC core
outrigger system was utilized whereby the core structure required very thick RC walls so as to meet the MCE level shear demands. VCDs were introduced to significantly reduce seismic demands and core wall thicknesses in the short direction which was very large when conventional flag walls (RCoutriggers) or BRBs were considered in preliminary design. The VCD system provided additional

703

704

CHAPTER 8 Case studies of tall buildings

benefits of performance enhancement under frequent wind storms as well as frequent earthquakes. Compared to the conventional design, the VCD outrigger design offered a more cost-effective as well as a more high-performing, safe, and resilient structure.

8.1.13.3 Structural system The lateral load resisting system of the building consists of an RC-coupled wall system in the EW direction and an RC core
outrigger system in the NS direction (Fig. 8.131). There are two outrigger levels along each of the NS direction core walls, one near two-thirds of the height of the building which spans three stories, where a VCD outrigger was used and the other at the top, which spans two stories, where a conventional RC flag wall outrigger was used (Fig. 8.136).

8.1.13.3.1 Building fundamental periods The fundamental periods of the building using the service level building model are (Fig. 8.132) Mode 1: 6.74 seconds (lateral NS); Mode 2: 5.59 (torsional); and Mode 3: 5.24 (lateral EW). The total seismic weight of the building was computed to be 1127 MN, while the seismic base shear coefficient was 0.059. The structural model of the building is shown in Fig. 8.133.

8.1.13.3.2 Damping strategy utilized A VCD damped outrigger system which was carefully designed to have a minimum impact on the functional requirements of the building (not occupying any sellable real estate and not requiring any monitoring or maintenance) was used. In addition, there is a second line of defense whereby the structural steel embeds connecting the modular VE damper panels to the structure are capacity designed

FIGURE 8.131 Connor Tower: (A) typical plan; (B) elevation A
A (EW direction); and (C) elevation B
B (NS direction).

8.1 Distributed Damping Systems Case Studies

FIGURE 8.132 Connor Tower: fundamental frequencies and mode shapes.

and detailed to include a yielding fuse mechanism. In an overload scenario the yielding fuse activates and limits the force imparted to the RC columns and outrigger wall and thus the integrity of the structure is preserved.

8.1.13.3.3 Additional damping provided by the damping system For service level wind loadings, VCDs provided an additional damping of 2% in the NS direction of the building to reduce drifts and accelerations. For rare MCE events the added damping from the VCDs reduced the structural demands.

8.1.13.3.4 Building cost versus damping cost The damping system was selected from a detailed financial assessment comparing BRBs, VCDs, and conventional flag walls. The damping cost was under 2% of the building cost and was offset by the reduced concrete (1097 m3) and reinforcement (1440 tons) throughout the building, an increased sellable floor area of the building (74.82 m2), and reduced construction time. The net benefit was estimated to be $1 million (including labor and increased sellable space) which was approximately 1.5% of the total construction cost. This net benefit would be much more

705

706

CHAPTER 8 Case studies of tall buildings

FIGURE 8.133 Connor Tower: 3D structural model.

substantial in other areas of the world such as North America or Europe due to increased construction cost and real estate values. Moreover, with the damping system, the life cycle cost of the building is expected to be smaller compared with a conventional RC system.

8.1.13.3.5 Building code The design was carried out in accordance with the International Building Code (IBC, 2012), ASCE 7-10 (ASCE, 2010), ACI 318-16 (ACI, 2016), ANSI/AISC 360 (AISC, 2016), and ANSI/AISC 341 (AISC, 2016). Since the building falls outside the code-prescribed limits on height, a performance-based seismic approach was used. For performance-based design, Pacific Earthquake Engineering Research Center (PEER Center) Tall Buildings Initiative (TBI, 2010)

8.1 Distributed Damping Systems Case Studies

guidelines, as well as the Los Angeles Tall Buildings Structural Design Council (LATBSDC, 2017) guidelines were used. Nonlinear modeling followed general approach outlined in ASCE 41-13 (ASCE, 2013) and ATC-114 (ATC, 2016) supplemented by latest component level test results. VCDs were designed according to Japanese Society for Seismic Isolation, JSSI Manual for Building Passive Control Technology (JSSI, 2015), and followed the appropriate damping device guidelines in ASCE 7-10 (ASCE, 2010) and American steel design codes—ANSI/ AISC 360 (ASCE, 2016) and ANSI/AISC 341 (ASCE, 2016).

8.1.13.3.6 Peer-reviewed project The building design including the damping system design was peer reviewed by MKA, Seattle.

8.1.13.3.7 Design forces Wind forces were determined based on standard wind tunnel testing (Fig. 8.134). For seismic loadings, a site-specific study was carried out. The historical seismicity of the site as well as the design response spectra for the service (SLE), DE, and MCER are shown in Fig. 8.135. For nonlinear time history analyses, seven ground motion pairs were selected and scaled per ASCE 7-2010.

8.1.13.3.8 Expected performance For the wind design, the interstory drift ratios under 1/10 year storms were less than 1/400 and the roof drift ratio was less than 1/500, which was achieved through the 2% added damping from the VCDs in the primary mode of vibration. The floor accelerations met the ISO criteria (ISO, 2007) for residential buildings for 1/1 year and 1/10 year storms. Structural members were designed for 1/700-year

FIGURE 8.134 Connor Tower: model of the building and its surroundings in wind tunnel.

707

708

CHAPTER 8 Case studies of tall buildings

FIGURE 8.135 Connor Tower: (A) historical seismicity in the area and (B) 5% damped design response spectra.

FIGURE 8.136 Connor Tower: MCER-level story demands from structural analysis: (A) NS direction shear forces and (B) bending moment about EW direction.

ultimate-level wind loadings. The seismic performance of the tower was more interesting because seismic forces mainly governed the design of the lateral system and the damping system. Fig. 8.136 shows the MCER-level building shear forces and bending moments with the damping system (with VCDs) compared to a conventional flag wall system (without the VCDs) and it is clear from this figure that the VCDs provided significant reductions in the forces. This trend is similar when comparing the VCD damped system to a diagonal BRB system, whereby the VCDs had reduced core demands. Finally, the building model was analyzed using nonlinear time history analysis using bounded properties of the VCDs. The results of the

8.1 Distributed Damping Systems Case Studies

Perform-3D model using upper bound and lower bound damper properties are shown in Fig. 8.137. Sample VCD hysteresis for the service as well as MCER-level ground motions are shown in Fig. 8.138 for a sample earthquake loading which demonstrate a well-defined force
deformation hysteresis for the respective earthquakes.

8.1.13.4 Damping overview 8.1.13.4.1 Damping strategy Based on the functional and performance requirements as well as a detailed cost
benefit analysis, a concentrated damping system, which consists of a VCD

FIGURE 8.137 Connor Tower: average response indicators from nonlinear time history analyses: (A) NS direction interstory drift, (B) EW direction interstory drift, and (C) coupling beam plastic rotations.

FIGURE 8.138 Connor Tower: average response indicators from nonlinear time history analyses: Typical VCD hysteresis under scaled M8.8 Chile, 2010 Subduction Earthquake (Municip. San Jose de Maipo Station): (A) SLE-level earthquake and (B) MCER-level earthquake.

709

710

CHAPTER 8 Case studies of tall buildings

FIGURE 8.139 Connor Tower: VCD schematics.

damped outrigger, was utilized at a single outrigger location. The layout of the dampers in the building is shown in Fig. 8.139A
C, a modular VE damper panel is shown in Fig. 8.139D, and 3D cross section of the VCDs is shown in Fig. 8.139E.

8.1.13.4.2 Damping type The VCDs consist of 3M ISD-111H VE material alternately bonded between steel plates. The damping material used herein is a fourth-generation 3M VE material developed for tall and super-tall building applications. The material is characterized by high stiffness and high damping as well as good durability. Two VE damper panels are bolted on two W-sections which are anchored into the RC structure forming the VCD. In order to avoid damage to the connecting structure in the event of a rare MCE, a lockup mechanism was designed such that VCDs lock up at 300% VE strain. For this design, the connecting steel elements are designed as shear-critical fuses that effectively dissipate energy under only large seismic demands (MCER-type loads). The steel connecting fuse elements can be inspected after a major seismic event and repaired or replaced if required. For all other events, such as strength level winds, the connecting elements are designed to remain linear elastic. In addition, a replaceable detail was designed for the connecting elements to further reduce downtime in the event of a strong earthquake and it will be decided by the design team whether it will be used at a later date. Capacity design principles were followed for the building as well as for the design of the damping system.

8.1 Distributed Damping Systems Case Studies

8.1.13.4.3 Structural and damping design Code and guidelines. As stated in Section 2.5, a performance-based design procedure per international standards was adopted. PEER-TBI (2010) and LATBSDC (2017) guidelines were used for the PBSD. The VCD design followed the damping device guidelines in ASCE 7-10 (ASCE, 2010) and JSSI (2015), as well as the performance-based design guidelines. The steel and connections were designed according to the American design codes—ANSI/AISC 360 (ASCE, 2016) and ANSI/AISC 341 (ASCE, 2016). Analysis modeling and software. Analysis and design was carried out using ETABS (for overall building), SAFE (for the foundation), Adapt Builder (for posttensioned slabs), CSiCol (for shear walls), spColumn (for columns), and Perform3D (for nonlinear time history analysis) (CSI, 2016). Linear static and nonlinear static (pushover) analyses were used for the preliminary design for wind loadings, while linear dynamic analyses (response spectrum analyses) were used for seismic design. Fast nonlinear analysis was used to determine the amount of added damping using a free vibration technique in ETABS. Damping performance for wind loadings was also verified using fast nonlinear analysis utilizing wind loading time histories generated from wind tunnel testing in ETABS. Final verification of the seismic design was carried out using nonlinear time history analysis under the suite of ground motions selected and scaled to represent the MCER-level shaking at the site. All analyses were carried out using state-of-the-art 3D models. For nonlinear analyses, VCDs were modeled using a spring element to simulate the connection stiffness in series with a generalized Maxwell model (GMM) which consists of one spring element in parallel with two Maxwell elements (spring and dashpot in series). The GMM is capable of simulating frequency dependency of the VE material and has been extensively validated using fullscale test results of the VCDs. Connecting steel elements were modeled using nonlinear link elements with a trilinear hysteresis in accordance with ASCE 4113 (ASCE, 2013) supplemented by latest test results. Design principles. Table 8.34 provides overall wind as well as seismic performance design objectives adopted for this project. These design objectives are Table 8.34 Connor Tower: Overall Wind and Seismic Performance Objectives Hazard

Level

Performance Objectives

Wind

Service (1- and 10-year return period) Strength (700 years) Service (43-year return period) DE (two-thirds of MCER effects) MCER (1% probability of collapse in 50 years)

Essentially elastic response; ISO (2007) occupant comfort criteria and drift criteria met Essentially elastic response Essentially elastic response

Wind Seismic Seismic Seismic

Life safety (equivalent to a code-designed building) Collapse prevention

711

712

CHAPTER 8 Case studies of tall buildings

consistent with the current state of practice of the PBSD in the west coast of the United States and wind design using ASCE 7-10 (ASCE, 2010). For damper design, a nonlinear modeling-based bounding analysis approach was used in accordance with ASCE 7-10 (ASCE, 2010) and PEER-TBI (2010). Design phase considerations. Uncertainties in damper properties were considered in all the design phases and all levels of wind and seismic hazards through a bounding analysis. The bounding values of the damper properties account for production variability, temperature variability, and mechanical property variability over a lifetime. Because heat is generated within the damping material due to vibrations, the temperature rise of the VE material was verified based on timehistory analysis using time histories generated from the wind tunnel testing as well as the ground motion time histories and was incorporated into the bounding analysis. For the slabs on the top of the VCDs, a special detail was used where the slab thickness was reduced such that the movement of the dampers could be maximized. Final deliverables of the project include drawings of the structural members and reinforcement details as well as the damper and lockup details and drawings as well as performance-based design report outlining the performance basis, modeling assumptions, as well as comparison of the demands with the acceptance criteria.

8.1.13.4.4 Architectural integration strategy The possibility of adopting a damped outrigger system was discussed among the architects, structural engineers, and damping consultants in very early in the design phase. This resulted in a minimum impact of the damped outrigger system on the architectural layout of the building. The location of the outrigger along the height of the building was largely dictated by the architectural requirements. In fact, the VCD damped system resulted in a superior architectural layout for the building with reduced wall thickness and reduced number of flag walls up to the height of the building allowing the developer to generate increased revenue as well as resulting in a more cost-effective structure.

8.1.13.4.5 MEP integration strategy The damped outrigger system provides space along the corridor with the core wall as well as near the column, which allow the passage of MEP systems without any problem.

8.1.13.4.6 Elevator and other mechanical devices integration strategy The damper configuration does not intervene with the elevators and other mechanical devices and thus no such integration was necessary.

8.1 Distributed Damping Systems Case Studies

8.1.13.4.7 Experimental tests The ISD-111H VE material as well as the damping system has been extensively tested for over 15 years at 3M, Nippon Steel and Sumikin Engineering, Tokyo Institute of Technology, University of California, Berkeley, and University of Toronto under a variety of loading protocols representative of wind as well as seismic loadings. Sample test setups and results are discussed in another case study of the 454 Yonge tower in Toronto (Section 8.1.9). Moreover, per standard practice, VE material and nondestructive full-scale uniaxial VE production tests are conducted on a represented number of dampers.

8.1.13.4.8 Monitoring system The VCD damped outrigger system does not require any inspection, maintenance, and tuning, so no monitoring systems is required. The durability of the 3M VE material has been validated based on more than 50 years of similar VE material types and more than 30 years of this particular VE material. In over 250 projects (more than 40,000 VE dampers) there has never been a single damper that has required replacement. As such, the ISD-111H VE material used in VCDs has shown and expected to not have any change in material properties over the lifetime of the building. All buildings in the Philippines, including this building, built above 50 m are monitored using accelerometers as per the Republic of the Philippines and the Department of Public Works and Highways standard to increase the understanding of tall buildings and detect damage after a severe seismic event.

8.1.13.4.9 Maintenance strategy As stated before, the VCDs do not require any maintenance or any specific care. The project team decided to use gypsum sheets the same as that required for a 2-hour fire rating for a steel beam.

8.1.13.4.10 Development process After the architectural design of the building was prepared, structural engineers developed a preliminary numerical model of the building. The wind tunnel consultant as well as the seismic hazard consultants was then engaged to carry out detailed site-specific studies. In the meantime, the damping consultants were involved to investigate the potential benefits of a damping system. The damping consultants used the seismic hazard as well as the wind hazard data to design the damping system, which was then finally validated based on detailed linear and nonlinear models of the buildings developed in collaboration with the structural engineers. The damper manufacturer was consulted by the damper consultants throughout the project.

713

714

CHAPTER 8 Case studies of tall buildings

8.1.13.5 Comparison of different design strategies 8.1.13.5.1 Structural options considered A coupled wall system was effective and efficient for the EW direction of the building and hence no other structural options were considered. For the NS direction of the building, the structure required either stiffness or damping to control drift because of the aspect ratio and flexibility of the core. An RC core outrigger system was developed to control frequent wind drifts. A conventional RC flag wall outrigger system design was possible; however, it was costly and affected the architectural layout of the building. This was because of large shear demands under rare earthquakes introduced to the core from the stiff outrigger system resulting in a very thick RC core wall and thick basement foundations. This was an expensive structural system and resulted in lost sellable space over the height of the building because of the size of structural members and large number of outrigger levels. A damping system in the outrigger was desirable to reduce the wind drifts, seismic demands, and the overall size of the structural core which was affecting the architectural layout of the building.

8.1.13.5.2 Damping solution considered Two alternative damping systems were considered: (1) VCD damped outrigger systems and (2) BRB outrigger system four stories in height at each outrigger location. A qualitative comparison of the VCD damped outrigger as well as the BRB outrigger system together with the conventional RC outrigger system is presented in Table 8.35. Qualitative benefits of VCD damped outrigger system are very clear from this table. Architecturally the BRBs were difficult to accommodate as the residential corridors would have to be reconfigured when the BRBs meet the core. Moreover, quantitatively, under strength wind loading the BRBs had large force demands and the BRBs could not be allowed to yield which resulted in an increased size of BRBs, further resulting in larger seismic demands because of the increased stiffness leading to the core wall shear demand increasing. Essentially these architectural limitations and large shear demands resulted in the design team preferring both the VCD configuration and the conventional core outrigger system over the BRBs.

8.1.13.5.3 Cost
benefit analysis Damping cost. The cost of damping system including testing, transportation, and erection was less than 2% of the cost of the building. If this building was built in North America it is estimated that the cost would be less than 0.4% of the construction cost. Design implications. Table 8.36 provides a preliminary cost
benefit analysis of the VCD damped outrigger system. As mentioned previously the design team elected to use the VCD system as opposed to a conventional RC outrigger system or a BRB system and thus only the cost of the VCD system is discussed here. Note that Table 8.35 provides a conservative estimate of the lowest possible

Table 8.35 Connor Tower: Comparison of Different Alternative Solutions Structural performance

Constructability and speed of construction

Architecture

Conventional RC Outrigger

VCD Outrigger

BRB Outrigger

Conventional RC outriggers provide stiffness; however, they do not add damping to the building. The stiffness is generally quite useful for reducing drifts.

VCDs are effective for all levels of earthquakes and all wind storms. The damping improves human comfort under frequent wind storms (those occurring every year or more) and frequent earthquakes (those occurring once a lifetime or more) compared to an undamped building. This is especially true for the people in the units at the top of the building.

For seismic loads, the conventional RC outrigger increases the shear forces in the building core.

The damping also delays the onset of damage compared to the undamped building for both rare wind storms and rare earthquakes protecting the developer’s investment. The steel fuse elements in the VCDs can provide effective capacity designed forces applied to the structure.

BRBs are effective for very large earthquakes and typhoons. BRBs do not provide any damping under smaller vibrations and therefore do not improve wind performance or seismic performance for more frequent events other than the beneficial effects of additional stiffness. In some cases BRBs sized for earthquakes must increase their capacity to not yield during wind storms. They do not add damping to the structure unless the steel core yields.

Under severe seismic loads, RC outriggers can suffer damage and can transfer even higher forcers to the core walls if not properly capacity designed. All of which must be accounted for in the design. Conventional RC outriggers are located only at few locations along the height of the building and thus they have a minimum impact on the constructability and speed of construction.

The conventional outrigger systems can have some minor impact on the architectural layout, especially passage of the MEP systems near flag wall to column connection.

Steel embeds to connect the dampers to the structure are constructed simply on-site as the outriggers are built. During the erection sequence a steel template is used to preserve the bolt alignment during construction. VE damper modular units are compact and are easily connected on a single floor with a small crew off the critical path of the structure construction. Through integration and design optimization with the structural engineers a VCD configuration that minimizes the number of outrigger floors and outrigger wall thickness can be developed for the VCD configuration.

The BRBs can reliably provide a maximum force applied to the columns and the outrigger walls.

BRBs are very large structural steel elements when applied as outrigger trusses. BRBs are required to be installed diagonally inbetween multiple floors using external cranes on the critical path of the structure. Connections and detailing can be relatively complicated in RC construction.

The BRBs can be located over multiple floors throughout the height of the structure, but at the times can impede corridor flow at the core levels.

716

CHAPTER 8 Case studies of tall buildings

Table 8.36 Connor Tower: Upfront Cost
Benefit Analysis Comparing VCD Structure to Conventional Flag Wall Construction Reduction in rebar (tons) Reduction in RC (m3) Increased sellable space (m2) Total savings Net savings

Quantity

Estimated Savings

1,440 1,097 74.82



$1.30 million $0.13 million $0.41 million $1.84 million $1 million

up-front cost savings. In Manila, RC construction is relatively inexpensive due to the relatively low cost of labor. Overall, the additional revenue for the developer was estimated due to the reduced shear wall thickness up to the building height and the number of flag walls and also due to resulting increased sellable space.

8.1.13.6 Lesson learned and recommendations 8.1.13.6.1 Difficulties in the design A design basis document considering bounding values of damper properties was prepared following the approach outlined in the performance-based design guides of PEER-TBI (2010) and LATBSDC (2010) guidelines by damper engineers. Since a performance-based design approach was to be used for even the conventional building, the inclusion of dampers in the building did not invoke any additional design requirements other than damping coefficients to be considered in the design and modeling. The ultimate shear capacity of the VCD was limited to the structural steel fuse elements which lead to a reliably designed capacity of the connecting columns and outrigger wall.

8.1.13.6.2 Design innovative solutions The innovative solution revolved around using a mature technology in a new configuration that is effective against both wind and seismic loadings and does not have any major impact on the architectural layout of the building. This was a straightforward design process.

8.1.13.6.3 Possible improvements No further improvements necessary for the damper design were considered in this project.

8.1.14 ALLIANZ TOWER, MILAN, ITALY 8.1.14.1 Project data The major building data (Fig. 8.140) are summarized as follows: • •

Year of completion: 2015 (topped out) Developer: CityLife S.p.A. (Italy)

8.1 Distributed Damping Systems Case Studies

FIGURE 8.140 Allianz Tower: building overview.

• • • • • • • • • •

Contractor: Colombo Costruzioni S.p.A. (Italy) Owner: Allianz S.p.A. Architectural designer: Arata Isozaki and Associates (Japan) and Andrea Maffeis Architects (Italy) Structural designer: Arup (Italy and International), ECSD S.r.l. (Italy), Studio Iorio S.r.l. (Italy) Damping supplier: FIP Industriale S.p.A. (Italy) Height: 202.20 m from the street level Interstory height: 3.9 m Stories: 50 above grade 1 3 below grade Gross area of the tower: Rectangular plan having dimensions L 5 24 m 3 61 m, total floor area 1464 m2, total surface 77.600 m2 Building function: Office building (private owner)

717

718

CHAPTER 8 Case studies of tall buildings

•

•

Structural material Concrete classes: Piles C32/40; foundation mat: C32/40 LH SCC (continuous massive pour); slabs: C40/50; cores: C40/50; columns: C70/85, C50/60, C40/50; top belt trusses: C65/75 fiber reinforced structural steel; composite columns; mid-height belt trusses; steel trusses; antenna; dampers: EN 10025-2 S355J2/K2 1 special elements Rebar: B450C Prestressing systems: DM 14/01/2008 1 European Technical Approvals Floor plan area: Rectangular plan having dimensions L 5 24 m 3 61 m, total floor area 1.464 m2

8.1.14.2 Introduction/history The tower is part of the CityLife in Milan which is a new residential, commercial, and business district partially still under construction. The Allianz tower is part of a complex of three different towers and its peculiar shape is called the “Straight one.” The tower is composed by 8 moduli of 6 floors each and the slightly curved fac¸ade on the outside provides a sensation of slight vibration of the building volume as it rises upward.

8.1.14.3 Structural system The main resisting system of the tower is a hybrid steel concrete structure having a total of six RC cores. Four of them are used as stairways/lift cores and two of them as smaller cores on the short sides of the building (Fig. 8.141). The other main components of the hybrid structural system are the two longitudinal steel belt trusses, located at mid-height, and the two prestressed concrete belt trusses located at the top. The belt trusses are also connecting the stairways cores. The lateral resisting system is enhanced with eight external VDDs located at the base of four steel struts connected to the building on the long side at a

FIGURE 8.141 Allianz Tower: structural floor plan.

8.1 Distributed Damping Systems Case Studies

one-third of the height. The steel elements are 40 m long (on the east side) and 60 m long (on the west side) (Figs. 8.142 and 8.143). The gravity system is completed with a series of composite steel
concrete and RC columns (Fig. 8.144) with diameters ranging between D 5 170 cm and D 5 65 cm. In total there are n (5 1 5) perimeter columns and n (4 1 4) inner columns. Finally, the floors consist of a ribbed structure with 20-cm thick RC toppings and 50-cm deep T-beams.

8.1.14.3.1 Building fundamental periods The building fundamental vibration modes are shown in Fig. 8.145. The first mode (T1  6 seconds) is mainly flexural along the long direction, the second

FIGURE 8.142 Allianz Tower: location of mid-height (steel) and top (prestressed concrete) belt trusses.

719

720

CHAPTER 8 Case studies of tall buildings

FIGURE 8.143 Allianz Tower: vertical section with materials.

mode (T2  5 seconds) is mainly flexural along the short direction, and the third one (T3  2.5 seconds) is mainly torsional.

8.1.14.3.2 Damping strategy utilized For this particular project a series of eight localized external viscous dampers were utilized to add damping to specific modes. Indeed, the main goal was to improve service life behavior in terms of comfort under wind loads and the additional goal was to improve the ULS behavior under LS (life safety) earthquake load.

8.1.14.3.3 Additional damping provided by the damping system In case the whole dissipative capacity of the dampers is exploited, due to prolonged exposure to the “design” wind values, and the dampers are considered to be fully effective (the design service temperature of the devices is attained), the

8.1 Distributed Damping Systems Case Studies

FIGURE 8.144 Allianz Tower: composite steel
concrete and RC column sections.

FIGURE 8.145 Allianz Tower: fundamental frequencies and mode shapes.

721

722

CHAPTER 8 Case studies of tall buildings

FIGURE 8.146 Allianz Tower: dynamic testing results on the entire tower equipped with viscous dampers.

nominal additional damping on the second and third mode is 9%. In case the dampers are supposed to lose 50% of their efficiency, the additional damping on the same modes is reduced to 6%. The inherent structural damping without dissipative devices is lower than 1% for all the vibration modes (Fig. 8.146).

8.1.14.3.4 Building cost versus damping cost To choose the right technological solution to implement in the tower, different strategies were studied to determine the most appropriate one to achieve the prerequired structural and architectural performance levels that could also provide the most advantageous economical solution. Giving the peculiarity of the project, every analyzed scheme was a prototype since no standard suitable solutions were available for the building. For this reason, numerical analyses were coupled with experimental tests on reduced-scale devices. The peculiar performance requirements of the dampers, i.e., to enhance comfort under wind excitation, and to provide additional damping at the LS earthquake limit state, required their valves to be able to operate and maintain efficiency at two very different velocities (one of which is one order of magnitude larger than the other). Based on these strict prerequirements, the technological solutions designed by different producers relied on different energy dissipation methods, different geometrical configuration and dimensions, and the use of different viscous fluids devices. Based on the preliminary analyses, the costs of the explored solutions were estimated to be ranging between 0.8% and 1.5% of the cost of the structural system, depending on the different technological solutions.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.147 Allianz Tower: lateral view and cross section of one of the four external steel struts.

In the Allianz Tower, due to architectural reasons, the dampers were located externally, at the base of the four steel struts that connect the long sides of the building to the ground (Fig. 8.147). These special hollow core structural steel elements, respectively, 40 and 60 m long, were specifically forged and installed with a high precision procedure, which represent a very significant additional cost, i.e., 8% of the total cost of the structural system (Fig. 8.148). However, the benefit provided by the damping system consists in a more efficient performance under the design wind loads, with a reduction of the maximum acceleration of up to 50%.

8.1.14.3.5 Building code Several building codes were utilized together with other, nonbinding, reference documents. Among all, the most important are: •

•

Codes (binding): • “Norme Tecniche per le costruzioni,” D.M. 14/01/2008 (NTC, 2008) (Italian Building Code, in Italian) • Eurocode 0, Eurocode 1 (CEN, 2010), Eurocode 2, Eurocode 3, Eurocode 8 (CEN, 2003) • CNR-DT 207/2008 (CNR, 2008) “Istruzioni per la valutazione delle azioni e degli effetti del vento sulle costruzioni” (Italian Code for the evaluation of wind action and effects, in Italian) • EN 15129:2009 (CEN, 2009) “Anti-seismic devices” Reference documents (nonbinding): • FEMA 356 (FEMA, 2000) “Prestandard and commentary for the seismic rehabilitation of buildings”

723

724

CHAPTER 8 Case studies of tall buildings

FIGURE 8.148 Allianz Tower: side view of the topped out building after the installation of the steel struts.

• •

FEMA 450 (BSSC, 2003) “NEHRP recommended provisions for seismic regulations for new buildings and other structures” ISO 6897:1984 (ISO, 1984) “Guidelines for the evaluation of the response of occupants of fixed structures, especially buildings and offshore structures, to low-frequency horizontal motion (0.063
1.0 Hz)”

8.1 Distributed Damping Systems Case Studies

8.1.14.3.6 Peer-reviewed project The peer-review process was carried out by two different firms: Ramboll (United Kingdom) and J&A Consultants (Italy).

8.1.14.4 Damping overview 8.1.14.4.1 Damping strategy Without additional damping, the expected maximum accelerations at the top floors exceeded the acceptability criteria set forth in ISO 6897:1984 (ISO, 1984) and were very close to those stated by CNR-DT 207:2008 (CNR, 2008). The high flexibility and slenderness of the Allianz Tower makes it more sensitive to windinduced accelerations, especially as for across-wind response and torsional modes. For this reason, in order to improve the performance in terms of comfort, additional damping devices were designed to be located in four external steel struts connected to the tower at level 11 (Figs. 8.149 and 8.150). The dampers are designed to mitigate both wind and earthquake effects, even if their contribution is not taken into account in the capacity verifications at ULS: in terms of capacity, the building is fully self-sufficient with limited damage expected to develop for the LS limit state.

FIGURE 8.149 Allianz Tower: design drawing of one of the eight dissipative devices and its connection to the strut.

725

726

CHAPTER 8 Case studies of tall buildings

FIGURE 8.150 Allianz Tower: damping devices during laboratory tests and after installation at the bottom end of one of the steel struts connecting them to the building.

8.1.14.4.2 Damping type N. 8 two-way FVDs were installed, two in each strut, designed and produced complying with EN 15129:2009 (CEN, 2009), having the following features: Design life: 50 years (with maintenance). Nonlinear constitutive law: F 5 Cvα, with α 5 1 for v # 6 mm/s; nonlinear behavior threshold (α{1) starts for 6 mm/s (20%) , v # 8 mm/s (115%). Damping constant: C 5 160 kN (mm/s)α/single device. Main vibration period: 2.5 seconds , T 5 T2, T3 , 5 seconds. Stroke: s 5 300 mm, including quasi-static displacements associated with longterm phenomena (concrete creep of the tower, differential settlements, temperature, average values of wind-induced displacements). Design external temperature: T 5 210 C/40 C. Maximum friction force: 100 kN. Theoretical maximum design allowable loss of effectiveness at design temperature: 50% (measured only 5% during wind tests on final devices). Design limit to transmissible force: Fmax 5 6 1500 kN/damper (10%/ 2 15%).

8.1 Distributed Damping Systems Case Studies

Service life design conditions: TR 5 1 year: vmax 5 6 mm/s, Fmax 5 6 960 kN/damper (linear range of behavior) TR 5 50 years: vmax 5 10 mm/s, Fmax 5 6 1500 kN/damper (nonlinear range of behavior) Design for fatigue: according EN 1991-1-4 (CEN, 2010) and EN 1993-1-9 (CEN, 2009) ULS (earthquake) design conditions (collapse prevention limit state): Maximum design force at CLS: 5 Fmax 5 6 1500 kN/damper Maximum velocity at CLS: v 5 120 mm/s Maximum displacement at CLS: s 5 1130 / 2160 mm The dampers have a two-way hydraulic circuit, with different valves for service life and ULS operation, due to the different velocities associated with the two loading conditions. The structural steel parts were designed with a focus on toughness and through-thickness properties and very reduced tolerances and gaps at each pin connection. A strict monitoring and testing protocol during the production, installation, and operation of each component was enforced.

8.1.14.4.3 Structural and damping design Code and guidelines. The reference documents listed were employed. In Italy, no specific rules for the design of external viscous dampers for wind effects mitigation are enforced in the Italian Building Code (NTC, 2008), as opposed to the specific rules set forth for design for earthquake mitigation. The installed dampers comply with EN 15129:2009 (CEN, 2009), with the use of a special mineral fluid to adequately lubricate the valves for service life operability. The experimental tests carried out according to EN 15129:2009 (CEN, 2009) were integrated with a wind test as specified in the FEMA Standards (FEMA, 2000), which was also used to evaluate the global expected damping values. The evaluation of the wind action and the related comfort analysis were carried out according to CNR-DT 207/2008 (CNR, 2008). Analysis modeling and software. When the structural vibration modes are uncoupled and the global damping is lower than 30%, equivalent linear analysis can be used for design. Alternatively, more refined types of analyses should be adopted, such as nonlinear time history or complex modal analysis. In the present case study, the use of linear equivalent analysis was allowed. Nevertheless, both the structure and the damping system were modeled with different degrees of refinements. The 3D finite element structural model including the dampers was validated and interpreted by means of comparison to the results yielded by a simplified cantilever model with uniformly distributed masses and equivalent stiffness. Viscous dampers were modeled both assuming a linear equivalent constitutive law (only for wind conditions) and by adopting a specific constitutive law to carry out nonlinear analyses. In the preliminary design phase, the effects of slightly VE behavior were computed in order to provide an upper limit for the

727

728

CHAPTER 8 Case studies of tall buildings

FIGURE 8.151 Allianz Tower: executive design phase graphs representing the performance of the Allianz Tower in terms of maximum acceleration under different wind loading conditions, in the original configuration without dampers versus in the configuration with the chosen additional dampers.

elastic constant K so that the deviation from the ideal purely viscous behavior would be negligible. For all of the analyses, the commercial structural analysis software MidasGen (Midas, 2017) was used: such analyses can be replicated with any commercial software of the same kind. Design principles. The comfort performance was required by the developer in terms of guaranteed additional damping for the first three vibration modes, related to which prerequired acceptability thresholds for the maximum accelerations were also defined. The maximum allowed acceleration at the top floor due to wind having a return period of 1 year was set to 5 cm/s2. This value corresponds to 50% of the expected acceleration in the same conditions for the building without additional dampers (Fig. 8.151). Design phase considerations. In the preliminary design phase, the use of a TMD was considered but quickly discarded due to budgetary and functional reasons, which led to adopting the solution of external dampers (Fig. 8.152). When entering the executive design phase, the performance levels required by the developer were enhanced: the use of additional dissipative devices to be installed in the top belt trusses was also considered at that point, as per a request of the developer.

8.1 Distributed Damping Systems Case Studies

FIGURE 8.152 Allianz Tower: preliminary analysis graph representing the along-wind acceleration response of the Allianz Tower for wind blowing in Y direction, estimated according to EC1 (CEN, 2010) for return period equal to 1, 5, and 10 years and structural damping equal to 1% and 5% (frequency of vibration equal to 0.198 Hz) and compared with various comfort criteria.

The analyses carried out in the executive design phase, though, highlighted that the benefits of installing additional devices in the top belt trusses would be negligible; on the other hand, increasing the dissipative capacity of the dampers located at the base of the struts would allow the enhanced performance levels to be guaranteed. In fact, most of the additional damping provided by the devices located at the bottom of the struts is active in mitigating the vibrations associated with the second and third modes, and mostly due to the across-wind response, thus improving the global behavior of the building in the transverse direction, which is by far the most critical, due to the high slenderness of the tower along that axis. On the other hand, adding dissipative capacity at the level of the top belt trusses would not achieve dramatic improvement of the global performance because of the high stiffness of the belt trusses, which reduces the relative velocity between the elements where the dampers would have been located, and

729

730

CHAPTER 8 Case studies of tall buildings

because most of the additional damping would have been effective only along the strong axis of the building, where the response is not critical. The final design of the dissipative devices consisted in a calculation report and a list of technical specifications to be complied with by the producers. A number of analyses were carried out on the finite element model of the tower including the dampers, assuming different conditions of operation: in this way, a comparative evaluation of the results was carried out to guarantee that the prerequired performance target would be met in any service life condition. All the data that the producers would need to design the dampers were thus provided. The dampers are CE marked. In Europe, the CE marking (a guarantee of “European Conformity”), which is the manufacturer’s declaration that the product meets the requirements of the applicable EC directives, makes the producer directly responsible for the declared specifications. All the design drawings, the technical calculation report, and the maintenance plan are also provided. These technical documents were drafted as the result of a close cooperation between the structural designers and the producer.

8.1.14.4.4 Architectural integration strategy The architectural concept by Arata Isozaki was inspired by the Futuristic Movement, focusing on the concept of the building machine. The steel struts, covered in golden paint, and the panoramic elevators aim at emphasizing the idea of exposing structural and functional elements and turning them into peculiar architectural features. The dampers are a fundamental part of this visible structural system. They are located in technical rooms at the top of the podium and at the first underground level.

8.1.14.4.5 MEP integration strategy There is no interaction between the damping system and the plants. The monitoring system for the struts and the dampers is self-sufficient and independent.

8.1.14.4.6 Elevator and other mechanical devices integration strategy There is no interaction between the damping system and the other mechanical devices. The monitoring system for the struts and the dampers is self-sufficient and independent.

8.1.14.4.7 Experimental tests A number of experimental tests were carried out on reduced-scale devices in order to experimentally determine the constitutive law and validate the performance in design wind conditions (Fig. 8.153). The results were very important to calibrate and validate the designed system. In the preliminary design phase, different products manufactured by a number of producers were tested to compare their performances. In the executive design phase, after choosing one producer, a set of quality certification tests were carried out on the devices, according to EN 1090:2012 (CEN, 2012) and EN 15129:2009 (CEN, 2009)—Cap 7.4, in addition

8.1 Distributed Damping Systems Case Studies

FIGURE 8.153 Allianz Tower: composite damping devices during laboratory tests and after installation at the bottom end of one of the steel struts connecting them to the building.

to energy efficiency tests according to FEMA 356 (FEMA, 2000) and 450 (BSCC, 2003). Cyclic wind tests were carried out: a continuous sinusoidal displacement time history (x(t) 5 X0sin(2πf)) was imposed for 2000 full cycles. The temperature developed in the dampers was continuously measured, together with the parameters needed to determine the constitutive law and the dissipated energy. The parameters used in design were: X0 5 maximum displacement amplitude associated with the fluctuating wind load, equal to a 6 5 mm for a return period of TR 5 1 year; f 5 0.2 Hz (the value corresponding to the second natural mode of the tower). The measured loss of efficiency at the end of the wind test was quantified in 5%, well below the acceptable design limit, initially defined as 50%. The experimental tests were supervised and certified by an independent third party.

8.1.14.4.8 Monitoring system Due to the complex structural system, a continuous monitoring system was designed and installed, acquiring data regarding the most significant structural members, including the dampers. The continuous acquisition of the wind at the top of the building (velocity and direction), of the horizontal accelerations of the top floor and of the displacement, temperature, and pressure of the eight dampers allow a wealth of data to be analyzed to monitor the performance of the dissipative system and the tower itself. Moreover, dynamic tests were carried out at the end of construction, both with and without dampers, so that the assumptions made in the design phase could be validated.

731

732

CHAPTER 8 Case studies of tall buildings

8.1.14.5 Comparison of different design strategies 8.1.14.5.1 Structural options considered Different options were analyzed to increase the global stiffness of the structural system in the longitudinal direction. The stiffening effect of two additional panoramic elevators, located at the NS ends of the building, was exploited. Moreover, the use of two belt trusses connecting the stairways cores contributed to further enhance the longitudinal stiffness, which eliminated the need for additional dampers effective in that direction. Different solutions for the belt trusses were studied: they were initially designed as steel trusses at both mid-height and at the top; subsequently, the use of prestressed concrete beams at both levels was proposed. The final chosen solution was hybrid, adopting a steel mid-height belt truss and a p.c. top belt truss.

8.1.14.5.2 Damping solution considered Two kinds of dissipative devices were thoroughly evaluated before making the final decision about the adopted system. One solution was the use of slightly viscoelastic, linear dissipative devices consisting of subelements connected in parallel to achieve the prerequired performance. This arrangement offered two benefits: reduced gaps at the connections and full efficiency in every operating condition guaranteed by the use of a special fluid used in aeronautical engineering. Nevertheless, it was incompatible with the architectural restraints, because each of the dampers would have occupied a volume of 6 m 3 6 m. The second solution was the use of linear viscous dampers with reduced dimensions. This system was adopted because of the compatibility with the architectural design and because the producer was able to enhance the initially required performance in a later phase with no significant increase in costs.

8.1.14.5.3 Cost
benefit analysis The cost
benefit analysis for the dampers was carried out by comparing the relative performance in terms of compliance with the design specifications. The factors deemed the most important were the inherent efficiency of the dampers, the interaction with the building at the design capacity limit, the robustness of the devices in case of misplacements, and/or inaccurate installation. Finally, the type of fluid was also considered. Damping cost. The cost of the chosen viscous dampers equals 0.8% of the cost of the structural system, whereas the cost of the four steel struts connecting the dampers to the building can be quantified as 8% of the cost of the structural parts.

8.1.14.6 Lesson learned and recommendations 8.1.14.6.1 Difficulties in the design No specific codes and prescriptions are provided in Italy for designing dissipative devices for comfort under wind excitation. Moreover, for tall buildings, it is not

8.2 Mass Damping Systems Case Studies

possible to use standard devices: every device is a prototype, in terms of both design and validation/testing. In the design phase, the use of equivalent linear analysis for wind conditions proved a very effective method to determine the basic prerequisites of the devices to be provided to the producers. For a tall building, the use of global dynamic tests for the validation of a dissipative system is difficult: when an operational modal analysis (OMA) is carried out, the environmental excitation is exploited to derive the dynamic properties of the system. This condition is not the one used to design the dissipative devices, so no direct comparison between the estimated frequency and damping values is possible for direct validation of the expected performance. This is especially true for environmental excitation analysis (OMA), but is also partly true for experimental modal analysis, i.e., when a known input excitation is provided by a rotating mass or an actuator. In fact, given the masses and the periods of such buildings, the induced vibration levels are always much lower than the design values. Still, coming up with a thorough dynamic testing and monitoring protocol for the whole building, coupled with extensive laboratory tests on the dissipative devices, allowed a more confident assessment of the efficiency of the dissipative system designed for the Allianz Tower.

8.1.14.6.2 Design innovative solutions The innovative features of the designed devices rely in the double system of valves, one for service life operability and the other one for ULS operability. Also, the two completely different constitutive laws of the dampers, one for service life, another for ULS.

8.1.14.6.3 Possible improvements In terms of comfort, the designed dissipative system guarantees enhanced performance levels with respect to those requested by the developer. Even at very low vibration levels, as measured in the OMA tests, the damping provided by the devices was about 300% larger than the inherent structural damping for Mode 2 and Mode 3. In terms of costs, the same performance could have been achieved by eliminating the external steel trusses and locating the dissipative devices in a different position. Nevertheless, the trusses represent a peculiar feature of the architectural concept, more than a structural necessity, so their added value remains in their architectural meaning.

8.2 MASS DAMPING SYSTEMS CASE STUDIES In this section, the case studies equipped with mass damping systems are reviewed. This refers to damping systems that dissipate energy through the differential movement of a large mass positioned in a discrete locations in a building (usually at the top).

733

734

CHAPTER 8 Case studies of tall buildings

8.2.1 CITICORP BUILDING, NEW YORK, NEW YORK CITY, UNITED STATES 8.2.1.1 Project data The major building data (Fig. 8.154) are summarized as follows: •

• • • • •

Year of completion: 1978 building 1977 TMD system Contractor: HRH Construction Company (United States) Owner: (Original) First National City Corp. (United States), (Since 2009) Boston Properties (United States) Architectural designer: Hugh Stubbins & Assoc. (United States) Project architect: Emery Roth & Sons (United States) Structural designer: Tower, Church, and TMD: LeMessurier (United States)

FIGURE 8.154 Citicorp Building: building overview.

8.2 Mass Damping Systems Case Studies

• • • •

• • •

• •

Retail base: Office of James Ruderman (United States) Damping supplier: Designed, built, installed by MTS System Corp. (United States) Testing laboratory: University of Western Ontario Boundary Layer Wind Tunnel Lab. (Canada) Height/interstory height • Height: 279 m above street level plus 3 basement stories • Interstory height: 3.9 m (typical floors) Stories: 59 above grade 1 3 below grade Gross area of the tower: 117,058 m2 Building function: • Tower: Offices • Base: Retail, offices, and church Structural material: Structural Steel Floor plan area: 2,244-m2 typical tower floor with 47.4-m side square plan

8.2.1.2 Introduction/history The top of this 59-story silvery aluminum and glass square tower is a 37.6-mhigh 45-degree sloped crown facing south. The bottom of the tower appears to stop 37.5 m above the street, except for a single column pair centered on each face and the central service core. This geometry was required to allow an existing Lutheran church to have its own free-standing identity beneath one corner of the tower. Early in the design phase, LeMessurier commissioned a wind tunnel study of the building from the Boundary Layer Wind Tunnel Lab at the University of Western Ontario. Their 1974 report showed that the tower as proposed would sway so as to make many of its occupants uncomfortable in wind storms. To solve this problem, the choices were either make a large increase in mass and lateral stiffness, or make a large increase in the building’s damping.

8.2.1.3 Structural system A steel supertruss in the plane of each exterior wall, in a repeated 8-story high chevron pattern, resists most of the lateral loads above 9-m deep transfer trusses which start at the 9th floor, 37.5 m above street level, and cantilever 20.4 m to each corner of the building. Below the transfer truss level, lateral loads are resisted by the diagonally braced central core (Fig. 8.155). All vertical loads in the wall planes are resisted by a single column pair at the center of each building face. Typical floor construction comprises 6.4-cm concrete on 7.6-cm steel deck on W21 steel beams at 3.8
3.9 m at the center; W24 steel girders spanning 13.1 m; and W30 diagonal steel girders spanning 16.4 m (Fig. 8.156). The structural steel weight an average of 107 kg/m2. Concrete foundations bear directly on bedrock. See Figs. 8.157
8.159 for photographs of the lateral structural system under construction.

735

736

CHAPTER 8 Case studies of tall buildings

FIGURE 8.155 Citicorp Building: typical structural elevations.

8.2.1.3.1 Building fundamental periods Originally calculated first mode of vibration periods were 6.7 seconds (NS), 7.0 seconds (EW), and 3.7 seconds (torsion). Mode shapes were nearly linear. Measured periods in 1986 were 6.96 seconds (NS) and 7.29 seconds (EW). Measured periods in 2005 were 7.04 seconds (NS) and 7.60 seconds (EW).

8.2.1.3.2 Damping strategy utilized The basic building damping was estimated to be 1.0% of critical. In order to reduce the wind-induced dynamic response accelerations to not more than 0.02020 milli-g for a 10-year storm, the damping had to be increased to 4.0% of critical. In building design, there was no precedent for adding this much damping, so LeMessurier proposed introducing a tuned mass damper (TMD) at the top of the building.

8.2 Mass Damping Systems Case Studies

FIGURE 8.156 Citicorp Building: typical framing floor.

8.2.1.3.3 Additional damping provided by the damping system The damping system was designed to provide added damping of 3.0% of critical, for a total of 4.0% in each major axis (NS and EW).

8.2.1.3.4 Building cost versus damping cost Roughly $150 million versus $1.5 million for the TMD system. The added cost of increasing the tower mass and lateral stiffness, if no added damping were provided, was estimated to be at least $5 million.

8.2.1.3.5 Building code New York City Building Code 1968 (NYCBC, 1968) (no seismic requirements.)

737

738

CHAPTER 8 Case studies of tall buildings

FIGURE 8.157 Citicorp Building: tower structural system under construction.

8.2.1.3.6 Peer-reviewed project The TMD system was reviewed by Simpson, Gumpertz, and Heger (United States). The building structure was reviewed by Les Robertson & Assoc. (United States) beginning in August 1978.

8.2.1.3.7 Design forces Wind loads: • •

NS: Base shear 5 18,754 kN, OT moment 5 2,914,673 kNm EW: Base shear 5 17,619 kN, OT moment 5 2,607,133 kNm

FIGURE 8.158 Citicorp Building: lower floors structure under construction.

FIGURE 8.159 Citicorp Building: upper floors structure under construction.

740

CHAPTER 8 Case studies of tall buildings

8.2.1.3.8 Expected performance Wind loads: • •

Peak accelerations of 20 milli-g peak with 0.10 a 10% annual probability Drift: H/600 for a uniform wind load of 0.957 kN/m2

8.2.1.4 Damping overview 8.2.1.4.1 Damping strategy A single biaxial TMD consists of a laterally sliding mass block near the top of the tower, at the 59th floor (243 m above street). This is above the highest occupied floor. The mass block was designed to have 2.0% of the generalized mass of the tower, about 0.5% of total tower mass, so as to provide a total of 4.0% damping. The TMD system was designed to act as a passive classical TMD (Fig. 8.160).

8.2.1.4.2 Damping type A 371,946 kg RC mass block, 2.7 m2 3 2.5 m high, is supported on twelve 0.56m diameter pressure-balanced oil bearings. The bearing sliding slide on an oil film on a large, specially leveled, grouted, and ground 1.91-cm steel plate, which is supported on a 30.5-cm thick RC slab on steel beams. The resulting horizontal friction coefficient is 0.003. The mass block is located on the geometric center of the building floor plan and is connected to the building structure by two linearmotion assemblies which provide pneumatic spring action as well as controlled

FIGURE 8.160 Citicorp building: upper floors structure under construction.

8.2 Mass Damping Systems Case Studies

FIGURE 8.161 Citicorp Building: TMD installed in building with linear-motion assemblies.

FIGURE 8.162 Citicorp Building: TMD hydraulic block snubbers.

hydraulic actuator force in each direction. This results in a passive biaxial TMD which has adjustable vibration periods in each direction, and a damping force (14% of critical) in each direction. The nonlinearities of the mechanical elements (springs and damping actuator) are compensated for by the servocontrol hardware and the active force capability of the servohydraulic actuators (Figs. 8.161 and 8.162). The desired TMD period is set to 1.02 times the first mode building period. The mass block is prevented from twisting about a vertical axis by a steel torque box antiyaw device, and will impact hydraulic snubbers against concrete curbs at floor level in the event of travel beyond its normal motion limit of 1.14 m. Sixteen snubbers are provided, each with a 0.3-m stroke. Accelerometers

741

742

CHAPTER 8 Case studies of tall buildings

on the floor provide continuous data to the TMD control system, which also includes accelerometers on the mass block and displacement transducers. The TMD remains stationary, on set-down blocks, until the lateral building acceleration exceeds 3 milli-g (0.003 g), at which point the hydraulic power system sends pressurized oil to the pressure-balanced bearings, lifting the mass block so it can move laterally. In case large motions cause the mass block snubbers to impact the curbs, the control system causes the TMD system to shut down and the mass block to return to its at-rest condition. Otherwise the control system causes shutdown at the end of a wind storm.

8.2.1.4.3 Structural and damping design Code and guidelines. Code: NYC Building Code 1968 (NYCBC, 1968) Guidelines: No specific guidelines were applicable. Analysis modeling and software. TMD parameter studies were conducted using a white noise forcing function applied to a two degree-of-freedom model with a large mass/spring/dashpot (building first mode, connected to the “ground”) and a small mass/spring/dashpot (the TMD mass block connected to the large mass via the spring and dashpot). The results were RMS values of horizontal displacement, which were converted to RMS accelerations. Some time history dynamic studies were made using a similar model but applying a modeled dynamic wind force, in order to provide estimates of maximum forces versus time for the system elements connecting the TMD mass block to the building. Similar time history analyses were made to validate acceptance testing, during which the control system, in test mode, placed the mass block and the building in motion and then switched to place the TMD system in passive TMD operational mode, which reduced the building motion displacements to zero. Design principles. The TMD system activates when the lateral building acceleration at the damper floor exceeds 3.0 milli-g. Criteria for TMD system shutdown include: • •

In a wind storm, if the mass block snubbers impact the curbs (i.e., when the mass block displacement from center exceeds 1.14 m). At the end of a wind storm, shut down when building accelerations are less than 1.0 milli-g for longer than 30 minutes.

Damping values: the building inherent damping is assumed to be 1.0% of critical. •

•

The TMD mass block damping is estimated to be 14% of critical. A smaller value would be somewhat more effective but would cause larger mass block displacements. The effective damping added to the building by TMD 3.0% of critical resulting in a total assumed damping of 4.0% of critical when the TMD is activated.

TMD periods: In each direction are 1.02 3 building period in each direction. (NS, EW). Only the first mode of building vibration was considered in each direction.

8.2 Mass Damping Systems Case Studies

Design phase considerations. During the preliminary design phase, several TMD configurations were investigated, each with a large mass, including hanging pendulum, hairpin shape steel “springs,” and torsional steel “springs.” LeMessurier found detailed design of such a system to be beyond the structural engineering domain, and in the realm of advanced mechanical engineering. David N. Wormley, a Mechanical Engineering Professor at MIT, was engaged as a consultant to help resolve issues such as: • • • •

The mass block-building interface had to include very low friction and absolutely safe gravity support; The “springs” had to be “tunable” (stiffness adjustable to provide a range of TMD periods) and had to be highly fatigue-resistant; A reliable “dashpot” system had to be provided and also must be adjustable; There had to be mass block overtravel restraints which were 100% safe, yet not produce excessive impact forces on the building.

In working through these issues with Professor Wormley, the decision was made to provide a detailed TMD performance specification, and with it solicit proposals from three or four suitable companies. During Design development MTS Systems Corp., in Minneapolis, MN, was awarded the contract to design and build most of the TMD components (except for such items as the mass block and concrete curbs). MTS worked closely with LeMessurier during the remainder of the design, manufacturing, and installation phases. Relevant structural, architectural, and MEP drawings were provided to MTS for coordination, and MTS provided preliminary and final design drawings for the TMD system, including detailed TMD
building interface drawings. MTS and LeMessurier also worked closely to develop acceptance testing procedures and acceptance criteria for the TMD system. Upon completion, MTS provided to the owner a detailed TMD system operations manual for the system.

8.2.1.4.4 Architectural integration strategy Since the TMD system was part of the original building design, its location on the 59th floor, a basically unfinished large room, was accounted for from the start. Wall locations were adjusted as needed to accomodate the TMD system design.

8.2.1.4.5 MEP integration strategy Lighting and ventilation for the TMD room were provided. Electric power and water specific to the TMD system were provided based on coordination between MTS and the building MEP engineers. Cooling water, up to 0.151m3 per minute, was required for the hydraulic power system (oil pumps). The TMD control system was linked to the building MEP central control system, to communicate current status information of the TMD system.

8.2.1.4.6 Elevator and other mechanical devices integration strategy No elevator service was available to the 59th floor TMD room. The building contractor provided crane service for building the cast-in-place concrete mass block, concrete curbs, and equipment pads, and for lifting TMD equipment into place.

743

744

CHAPTER 8 Case studies of tall buildings

8.2.1.4.7 Experimental tests Tests of an aeroelastic model of the building and its surroundings were conducted by the wind tunnel laboratory. Special testing was required and provided by MTS for the overtravel snubbers. Each of the 16 snubbers was subjected to a proof test in which a railroad boxcar was rolled into the snubber at a specified velocity. TMD system acceptance tests were conducted after the building periods were measured and the TMD system was complete. The TMD was put into a mode where the controls forced the mass block to push the building into motion at its first mode of vibration, to an approximate acceleration of 10 milli-g, then the controls were switched to place the TMD into passive response mode, which caused the building lateral motion to decay to zero.

8.2.1.4.8 Monitoring system Accelerometers mounted at the building 59th floor continuously monitor building lateral movement. When the building lateral acceleration reaches the TMD startup threshold, the system automatically records all TMD instrument values until the system shuts down.

8.2.1.4.9 Maintenance strategy MTS Systems Corp. has provided periodic maintenance on a contract basis following the initial warranty period. MTS provided a TMD system operating manual to the Owner and helped train building maintenance staff.

8.2.1.4.10 Development process The use of a TMD system was recommended as a design option by the Wind Tunnel Lab and tested by them in an aeroelastic test on a model of the building. Wind tunnel tests were conducted on the building model with and without the TMD, at LeMessurier’s request. Based on the Wind Tunnel Lab’s report and structural engineering studies, LeMessurier proposed use of the TMD system as the best and most economical option to make the building wind response more comfortable for its occupants. LeMessurier, with the help of Stubbins and Associates, convinced the owner to proceed with this unprecedented solution.

8.2.1.5 Comparison of different design strategies 8.2.1.5.1 Structural options considered In order to achieve the same effect on building wind response as the TMD system, the structural engineers would have had to roughly double both the mass and lateral stiffness of the tower, adding 25,401 metric ton more concrete and steel. This would have cost at least 3 times as much as the TMD, in structure alone. It also would have caused an increase in story height and overall building height, with associated cost increases to cladding, mechanical, electrical, plumbing and vertical transporation.

8.2 Mass Damping Systems Case Studies

8.2.1.5.2 Damping solution considered In the preliminary design phase (see Section 8.2.1.4.3), LeMessurier considered distributed VE dampers similar to those used on the World Trade Center twin towers, but did not select that option, since the building floor framing system did not result in sufficient floor beam end rotations. Other TMD options considered are described previously in Section 8.2.1.4.3.

8.2.1.5.3 Cost
benefit analysis The TMD damping system was considered and selected in the conceptual design phase for cost savings and the probable reliability of achieving the added damping required.

8.2.1.6 Lesson learned and recommendations 8.2.1.6.1 Difficulties in the design Additional advanced analyses were required to design the building with a TMD system as compared to a building without a TMD system. The resulting design proved to be a highly effective method for achieving occupant comfort criteria while minimizing structural cost. Working outside the realm of ordinary structural engineering practice, and advancing the state of the art in the process, was therefore deemed to be worth the cost and effort.

8.2.1.6.2 Design innovative solutions The Citicorp Tower TMD system marked the first time that a TMD system was incorporated into the design of a tall building as a part of the original design concept. It was also the first design of a biaxial TMD. It, and was first use of a pair of opposing pneumatic cylinders as “springs.”

8.2.1.6.3 Possible improvements TMD with a smaller mass block and active control of the hydraulic actuators would have been an improvement; however, the reliability of computer control systems at the time would have been a cause for safety concerns. It would have been better for maintenance and replacement of large equipment elements for the TMD room to be served by a building freight elevator. It would be good to have wind measuring instruments (e.g., anemometers) mounted at the top of the tower; therefore both building and TMD responses to wind storms could be correlated.

8.2.2 JOHN HANCOCK TOWER, BOSTON, MASSACHUSETTS, UNITED STATES 8.2.2.1 Project data The major building data (Fig. 8.163) are summarized as follows: •

Year of completion 1976 Building 1977 TMD system

745

746

CHAPTER 8 Case studies of tall buildings

FIGURE 8.163 John Hancock Tower: building overview (copyright Anton Grassl/Esto).

• • • •

• • • • • • • • •

Contractor: Gilbane Building Co. (United States) Owner: (Original) John Hancock Mutual Life Insurance Co. (United States) (Since 2015) Boston Properties (United States) Architectural designer: I.M. Pei & Partners (United States) Structural designer TMD: LeMessurier (United States) Building: Office of James Ruderman (United States) Damping supplier: Designed, Built, Installed by MTS System Corp. (United States) Testing laboratory: University of Western Ontario Boundary Layer Wind Tunnel Lab (Canada) Height: 241 m above street level plus 2 basement stories Interstory height: 3.8 m Stories: 62 above grade 1 3 below grade Gross area of the tower: 167,225 m2 Building function: Offices Structural material: Structural Steel Floor plan area: 2,706 m2 typical tower floor; parallelogram in plan, 31.2 m 3 87.6 m to slab edges

8.2.2.2 Introduction/history Upon its completion in 1976, the 241-m tall John Hancock Tower was the tallest building in Boston, and it remains so till the publication of this case study (2018). Its completion was scheduled for 1972, but in November 1972 story-high panes of exterior glass began popping out and falling to the street. This continued for

8.2 Mass Damping Systems Case Studies

some time, and all the exterior glass on the building was redesigned and replaced (10,344 panes). The architect engaged two renowned experts (Prof. Bruno Thurlimann of Swiss Federal Institute of Technology and Prof. Alan Davenport of the Boundary Layer Wind Tunnel, University of Western Ontario) to investigate the tower in relation to effects on the windows from wind-induced structural displacements. Their study used state-of-the-art wind tunnel technology including probability curves of wind velocity and direction for the site and an aeroelastic building model which included the first three vibration modes calculated by Prof. Thurlimann. The building wind-induced displacements were found not to be a cause of the glass failures, but there were two alarming results of the extensive structural review by Prof. Thurlimann: (1) In the event of a 100-year recurrence wind storm the tower would have a safety factor near 1.0 in the longitudinal direction, and (2) for wind storms having a recurrence rate of even once a year the top of building lateral accelerations would reach 15
22 milli-g, which would make building occupants quite uncomfortable. Values for a 10-year recurrence would be 17
33 milli-g. The larger value is for the worst corner of the floor plan, where the east
west sway mode and the twist mode act together since their vibration periods are somewhat close. The owner and architect in early 1975 directed that the tower be strengthened extensively in the longitudinal (NS) direction. The original building had been designed for the Boston Building Code (BBC, 1970) wind loads, but was designed in both principal directions for a horizontal drift ratio of height/500 for a uniform lateral load of 957 N/m2. This was common New York City engineering practice at the time. The problem was the effective lateral stiffness is reduced greatly in the longitudinal direction by P-Delta effects. This was considered an insignificant “second-order effect” in 1970, but has been shown since to be significant for some buildings, especially those having an elongated floor plan. The architect engaged LeMessurier in early 1975 as consultants to solve the wind motion human comfort problem since LeMessurier was the designer of the building and the TMD system in the Citicorp Center Tower in New York City (see Section 8.2.1). The firm accepted the challenge of designing a TMD system as a retrofit in an already constructed building. The mass blocks and their supporting structure had to fit within the height limits of an upper office floor, since spaces closer to the roof were fully committed to a public observatory and an MEP equipment level. The TMD system had to increase building damping in both the transverse (EW) vibration mode and the twist (about vertical axis) vibration mode. The estimated building vibration periods were 7.0 seconds EW and 6.6 seconds twist. The objective for occupant comfort was to not exceed peak lateral accelerations, at any point of the top occupied floor, of 20 milli-g for a recurrence interval of 10 years. The building’s structural engineer surveyed the actual gravity loadings of the existing mechanical equipment at floor 61 and above, and found them to be much less than the original specified live loads, and advised LeMessurier how much added gravity load could be safely imposed on the existing columns. LeMessurier conceived a preliminary design of a dual TMD system having two 272,155 kg mass blocks located at 67.5 m apart at 219 m above the street. The design would add damping in both transverse and twist vibration modes.

747

748

CHAPTER 8 Case studies of tall buildings

A detailed performance specification was prepared in July 1975, and MTS Systems Corp. was selected to design and provide most elements of the dual TMD system. Excluded were mass blocks and, the structural support slab, and dunnage framing, which were designed by LeMessurier and built by the building contractor.

8.2.2.3 Structural system Structural steel framing, with lateral load resistance provided by: •

•

In the longitudinal direction (NS): two perimeter moment-resisting frames each consisting of nine bays of 8.7-m span W shape beams and W14 or builtup columns (Figs. 8.164 and 8.165). Note: heavy diagonal bracing in vertical plane on two column lines at long sides of the core was added in 1976
1977. In the transverse direction (EW): eight lines of W shape beams across the full width of the building, with bays 6.5, 3.6, 10.1, 3.6, and 6.5 m. The central bay is a diagonally braced frame, and the other four bays are moment-framed.

The structural height-to-width ratio is 240.7/30.3 5 7.94. Typical floors have 7.6 cm stone concrete on 7.6 cm cellular decking on W21 steel beams spaced at 3.2
3.6 m c/c, spanning 8.7 m. Foundations consist of steel H-piles driven to bedrock, with RC pile caps and a pressure slab.

8.2.2.3.1 Building fundamental periods Building fundamental periods: • • •

First mode: TNS 5 Under 11 seconds (final stiffened building) Second mode: TEW 5 7.36 seconds (7.44 seconds, measured in 1984) Third mode: TROT 5 6.67 seconds (6.68 seconds, measured in 1984)

8.2.2.3.2 Damping strategy utilized The basic building damping was measured to be: • •

East
west mode: 0.90% of critical (0.92% in 1984) Twist mode: 2.80% of critical (2.08% in 1984)

FIGURE 8.164 John Hancock Tower: typical framing plan.

8.2 Mass Damping Systems Case Studies

FIGURE 8.165 John Hancock Tower: DTMD supporting flooring plan.

In order to reduce wind-induced peak lateral accelerations to not more than 20 milli-g for a 10-year recurrence storm, the east
west damping would be increased to about 3.5% of critical and the twist mode damping would be increased to about 4.2% of critical. The twin 272,155 kg mass blocks provide a ratio of 1.4% of the EW mode generalized building mass and 2.1% of the twist mode generalized building mass.

8.2.2.3.3 Additional damping provided by the damping system The Dual TMD (DTMD) system was designed to provide added damping of about 2.6% of critical in the EW mode, and about 1.4 % critical in the twist mode, for totals of 3.5% in the EW direction and 4.2% in the twist mode).

8.2.2.3.4 Building cost versus damping cost Roughly $144 million versus $3.5
4 million for the DTMD system.

8.2.2.3.5 Building code Boston Building Code 1970 (BBC, 1970).

8.2.2.3.6 Peer-reviewed project The DTMD system was reviewed by Hansen, Holley and Biggs (United States). The building structure was reviewed by LeMessurier (United States). Comprehensive wind effects were reviewed by B. Thurlimann (Switzerland), and A. Davenport University of Western Ontario (Canada).

749

750

CHAPTER 8 Case studies of tall buildings

8.2.2.3.7 Design forces Wind loads: • •

NS: Base shear 5 12,866 kN, OT momentum 5 1,755,545 kNm. EW: Base shear 5 39,006 kN, OT momentum 5 5,700,171 kNm.

8.2.2.3.8 Expected performance Wind loads: • •

Peak accelerations of 20 milli-g at the worst corner of the top occupied floor with a 10% annual probability. Drift (EW): H/500 for a uniform wind load of 0.958 kN/m2.

8.2.2.4 Damping overview 8.2.2.4.1 Damping strategy A DTMD (Figs. 8.166 and 8.167) system is used, consisting of two laterally sliding mass blocks (each in only the EW direction) near the top of the tower (219 m above street). The DTMD was designed to increase the building effective damping in both the transverse (EW) mode and the twist mode. The DTMD system was designed to act as a modified passive classical TMD in order to meet the

FIGURE 8.166 John Hancock Tower: DTMD schematic diagram.

8.2 Mass Damping Systems Case Studies

FIGURE 8.167 John Hancock Tower: DTMD installed in building with a pressurized pneumatic (nitrogen) “spring,” a servo-controlled hydraulic actuator, and snubbers.

occupant comfort criterion. The building structure is designed to resist wind loads without participation from the damper system.

8.2.2.4.2 Damping type A modified classical TMD system was modified with two 272,155 kg mass blocks, near opposite ends of an upper office floor at 33.8 m from the EW axis of the tower, and centered 0.3 m from the NS axis. Each mass block is connected to the building with a pressurized pneumatic (nitrogen) “spring” and a servocontrolled hydraulic actuator (Figs. 8.168
8.169). The spring-mass system was specified to be adjustable over a range of 6.6
9.0 seconds vibration periods. Each mass block is constructed of a welded steel coffer box, filled with lead blocks, plates, and pellets (about 253,105 kg of lead). The box, with dimensions of 5.2 m 3 5.0 m 3 1.0 m, is supported on 40.6
55.9 cm diameter pressurebalanced oil bearings sliding on an oil film on four specially leveled and polished stainless steel plates. The resulting horizontal friction coefficient is 0.003. In order to avoid overloading the existing floor structure, all vertical and lateral loads from each TMD are transferred to the steel building columns via a specially designed platform consisting of a concrete slab on a 4.8 m 3 9.2 m 3 2.5 cm thick steel plate, bearing on ten W36 steel “dunnage” beams connected to five building columns. The mass block maximum displacement in each direction (EW) is 1.8 m before the block engages hydraulic snubbers which allow an additional 0.23 m displacement. Guide tracks are provided to prevent NS movement of the mass block. The DTMD control system provides for tuning the damper vibration period (in combination with the precharge pressure in the pneumatic “springs”), and for

751

752

CHAPTER 8 Case studies of tall buildings

FIGURE 8.168 John Hancock Tower: DTMD pressure-balanced oil bearings.

FIGURE 8.169 John Hancock Tower: pressurized pneumatic (nitrogen) “spring” and servo-controlled hydraulic.

a three-stage nonlinear TMD damping force from the hydraulic actuator. The three stages are: (1) For mass block displacements less than 1.5 m, provide a damping force corresponding to 7.0% of critical. (2) For any mass block displacement exceeding 1.5 m (except if it exceeds 1.5 m twice within an interval of 100 seconds) provide a damping force corresponding to 14.0% critical. (3) For any displacement exceeding 1.5 m or more within 100 seconds, provide a

8.2 Mass Damping Systems Case Studies

damping force as per stage 2 plus an additional damping force of 0.24 x(block displacement)2x (block velocity). The damping force does not to step in reverse order, such as stage 2 to 1, or stage 3 to 2. The nonlinearities of the mechanical elements (springs and damping actuator) are compensated for by the servocontrol hardware and the active force capability of the servohydraulic actuators. There are eight snubbers (shock absorbers), four at each end of each mass block, with specified properties, to control the mass blocks in the event of block displacements exceeding 1.8 m. The snubbers are specified to reset in less than 3 seconds. The desired DTMD vibration period is set to a value slightly larger than the building EW period. Accelerometers mounted in the TMD rooms provide continuous data to the DTMD control system, which also includes accelerometers on the mass blocks, and displacement transducers. The mass blocks remain stationary, on set-down blocks, until lateral building acceleration exceeds 3.0 milli-g, when the hydraulic power system starts up and sends pressurized oil to the pressure-balanced bearings, lifting the mass blocks so they can move laterally. In case of large motions cause the mass block snubbers to reach 75% of their total stroke, the control system shuts down the DTMD system and forces the mass blocks to return to their normal centered at-rest positions. Otherwise the DTMD system remains operational until shutdown at the end of a wind storm, specified as accelerations less than 1.0 milli-g for a period of 30 minutes. Each damper mass block has a self-test, or exercise, mode of operation, to be performed at scheduled intervals, as often as weekly.

8.2.2.4.3 Structural and damping design Code and guidelines. Code: Boston Building Code 1970 (BBC, 1970). Guideline: no specific guidelines were applicable. Analysis modeling and software. TMD parameter studies were conducted using a white noise forcing function applied to a two degree-of-freedom model with a large mass/spring/dashpot (building EW mode and, separately, building twist mode, connected to the “ground”) and a small mass/spring/dashpot (the TMD mass blocks connected to the large mass via the TMD springs and dashpots). The results were RMS values of horizontal displacement, which were converted to RMS accelerations. Some time history dynamic studies were made using a similar model, but applying a modeled dynamic wind force, in order to provide estimates of maximum forces versus time for the system elements connecting the TMD mass blocks to the building, in each mode of vibration. Some time history studies were also made using the second and third stage damping force equations for the large displacement conditions of the mass blocks impacting the snubbers. Time history studies included those made to validate acceptance testing, during which the control system, in test mode, placed the mass blocks and the building in motion and then switched to place the DTMD system in passive-DTMD operational mode, which then reduced the building motion displacements to zero. This was done for both EW and twist modes, during acceptance testing.

753

754

CHAPTER 8 Case studies of tall buildings

Design principles. Performance levels considered: The DTMD system activates when lateral building acceleration at either damper room exceeds 3.0 milli-g. Criteria for TMD system shutdown include: •

•

In a wind storm, if the mass block snubbers impact the fixed snubbers to 0.6 maximum stroke (i.e., when either mass block displacement from center exceeds 2.0 m). At the end of a wind storm, when building accelerations are less than 1.0 milli-g for longer than 30 minutes. Damping values:

•

The building inherent damping is assumed to be 0.9% of critical in the EW direction, and 2.8% of critical in the twist mode. In stage 1, the DTMD mass block damping is estimated to be 7.0% of critical, which is near optimal. • In stage 2, which occurs if either mass block displacement reaches 1.5 m, the damping increases to 14.0% of critical. • In stage 3, for larger mass block displacements, the damping increases more in a third stage (see Section 8.2.2.4.2). • The effective damping added to the building by the DTMD system is about 2.6% of critical in the EW mode, and about 1.4% of critical in the Twist mode, for totals of 3.5% in the EW direction and 4.2% in the twist mode. Design phase considerations. The owner requested design and installation of the TMD system to proceed as quickly as possible, since the building was to be occupied in late 1976. LeMessurier and its consultant Prof. David N. Wormley, a Professor of Mechanical Engineering at MIT, had been working together on the design phase of a biaxial TMD system for the Citicorp Center Tower in New York City (see Section 8.2.1). The John Hancock Tower DTMD system had to meet quite different requirements, mainly that it had to address not only transverse sway but torsional sway motion. It also had to be retrofit into an existing building which would be occupied during the installation. Structurally it would be desirable to have two-mass blocks spaced as far from the building center as possible, and with as much mass as possible. The owner did not want to sacrifice any more rentable floor area than absolutely necessary, and wanted to avoid any costly and disruptive reinforcement of existing steel columns below the damper locations if possible. LeMessurier believed that a DTMD System comprised of two mass blocks sliding only in the E-W direction could be provided by MTS Systems Corp., from Minneapolis, MN based on the experience with the Citicorp TMD system. Following intensive feasibility studies, a detailed DTMD performance specification was provided to MTS Systems Corp., to solicit their proposal. The design called for two 272,155 kg mass blocks (built of steel coffer boxes filled with lead) located 67.5 m apart on an office floor about 219 m above street level. The mass blocks and associated special dunnage steel platform were designed to prevent overloading the existing building columns.

8.2 Mass Damping Systems Case Studies

MTS Systems Corp. was awarded the contract to design and build most of the DTMD components (except for such items as the mass blocks, the structural support dunnage platform, fire protection, room acoustical walls, etc.) MTS worked closely with LeMessurier during the design, manufacturing, and installation phases. Relevant structural, architectural, and MEP drawings were provided to MTS for coordination, and MTS provided preliminary and final design drawings for the DTMD system, including detailed DTMD
building interface drawings. MTS and LeMessurier also worked closely to develop acceptance testing procedures and acceptance criteria for the DTMD System. Upon completion, MTS provided to the owner a detailed DTMD system operations manual for the system.

8.2.2.4.4 Architectural integration strategy LeMessurier worked with I.M. Pei & Partners, the Office of James Ruderman, and MTS Systems Corp. to establish exact DTMD room locations and dimensions during the preliminary design phase. The objectives were to locate the mass blocks as far apart as possible, avoid having to reinforce existing steel columns below damper rooms, and to limit size of the damper rooms. Each damper room needed to accommodate the necessary sliding mass block, the spring and actuator and snubbers, the hydraulic power supply, the control console, the support platform and its dunnage steel framing, and the “aisle” spaces needed to provide clearance for maintenance. The final size of each room was about 10.7 m 3 18.3 m. Other issues of concern to the architect and owner included: 1. Fire protection, in view of a significant quantity of hydraulic oil both in reservoirs, in piping and exposed. LeMessurier hired Firepro Inc. to design a Halon fire suppression system for each damper room. 2. Noise and vibration due to proximity of oil pumps in the TMD hydraulic power supply to the occupied office spaces above, below, and on three sides of each damper room. LeMessurier engaged acoustic consultant William Cavanaugh, who worked with MTS Systems Corp. and the architect to prevent or resolve any noise/vibration problems.

8.2.2.4.5 MEP integration strategy Lighting and air-conditioning ventilation for each damper room were supplied. Electric power and water specific to the DTMD system were provided based on the coordination between MTS and the building MEP engineers. Cooling water up to 0.151 m3 per minute was required for the hydraulic power system (oil pumps) in each damper room. The DTMD control system was linked to the building MEP central control system, to provide performance monitoring and safety checks of the DTMD system.

8.2.2.4.6 Elevator and other mechanical devices integration strategy Freight elevator service was available. Since the building was already fully enclosed and partly occupied, all DTMD elements and supporting platform elements had to be limited in length to 3 m, so as to fit within the elevator, and

755

756

CHAPTER 8 Case studies of tall buildings

limited in weight to under 3,175 kg due to elevator capacity. The steel dunnage beams and mass block steel coffer box plates had to be spliced with fullpenetration welds and fully inspected, after transport to their final locations. The lead blocks, plates, and pellets were brought to the damper floor and then transported to the damper rooms by conveyor belt for efficiency, and to avoid overloading the building floor. The Contractor, Gilbane Building Co., coordinated the transport work.

8.2.2.4.7 Experimental tests Special testing was required and provided by MTS for the overtravel snubbers. Each of the snubbers was subjected to a proof test in which a railroad boxcar was rolled into the snubber at a specified velocity. DTMD system acceptance tests were conducted after the building periods were measured and the DTMD system was complete. The DTMD was put into a mode where the controls forced the mass blocks to push the building into motion in its EW sway mode of vibration, to an approximate acceleration of 7.0 milli-g. Then the controls were then switched to place the DTMD into passive response mode, which caused the building lateral motion to decay to zero. This testing process was repeated for the building twist mode of vibration. In a third test, the north end damper pushed the building into its twist mode with accelerations of 6.5 milli-g, equal and opposite at north and south ends, then continued the north damper in active mode while switching the south damper to passive response mode. The south end accelerations were reduced by 45%.

8.2.2.4.8 Monitoring system Accelerometers mounted on the building structure at each damper room continuously monitor building lateral movement. When the building lateral acceleration reaches the DTMD start-up threshold, the system automatically records all DTMD instrument values until the DTMD shuts down.

8.2.2.4.9 Maintenance strategy MTS Systems Corp. has provided periodic maintenance on a contract basis following the initial warranty period. MTS provided DTMD system operating manual to the Owner and helped train building maintenance staff.

8.2.2.4.10 Development process The use of a DTMD system was recommended by the Wind Tunnel Lab and tested by them in an aeroelastic model of the building. Wind tunnel tests were conducted on the building model with and without the DTMD at LeMessurier’s request. Based on the Wind Tunnel Lab’s report and structural engineering parametric studies, LeMessurier proposed the use of the DTMD System as the best and most economical option to make the building wind response acceptably comfortable for its occupants. LeMessurier with the help of I.M. Pei and Partners, convinced the Owner to proceed with this unprecedented solution.

8.2 Mass Damping Systems Case Studies

8.2.2.5 Comparison of different design strategies 8.2.2.5.1 Structural options considered Since the building was about 90% complete and fully enclosed, an option to add significant mass and stiffness in both EW sway mode and the twist mode was neither architecturally acceptable nor economically feasible. The only solution to the building wind motion problem was to add a DTMD System near the top of the building.

8.2.2.5.2 Damping solution considered Only one solution was considered, as it was the only one that would add sufficient damping in both the EW and twist modes of vibration. The use of concrete mass blocks was not feasible, since the blocks would be too large in height and width to fit within existing story height and also to clear existing columns. Therefore lead-filled steel coffer boxes were used.

8.2.2.5.3 Cost
benefit analysis Such an analysis was not applicable, since there was no reasonable alternative.

8.2.2.6 Lesson learned and recommendations 8.2.2.6.1 Difficulties in the design The architectural constraints on room size and quite limited overall height of the damper mass block and its support structure contributed to the design complexity. The damper support structure had to be carefully planned in order to impose gravity loads on existing building columns within the added load parameters set by the building structural engineer.

8.2.2.6.2 Design innovative solutions The John Hancock Tower, Boston, was the first tall building, by a matter of months, to incorporate an operational TMD system, and the first to have a TMD system as a retrofit to an already occupied building. This was also the first TMD system to have twin mass blocks, at opposite ends of a floor, reducing wind dynamic response in both transverse and twist vibration modes. It was the first actual use of a pair of opposing pneumatic cylinders as “springs.” Note: similar pneumatic springs were used in the Citicorp Center TMD System, completed about 7 months later in 1977.

8.2.2.6.3 Possible improvements TMD with smaller mass blocks and active control of the hydraulic actuators would have been an improvement; however, the reliability of computer control systems at the time would have been a cause for safety concerns.

757

758

CHAPTER 8 Case studies of tall buildings

8.2.3 TAIPEI 101, TAIPEI, TAIWAN 8.2.3.1 Project data The major building data (Fig. 8.170) are summarized as follows: • • • • • • • •

Year of completion: 2004 Developer/contractor: Taipei Financial Center Corporation (Taiwan) Architectural designer: C.Y. Lee & Partners, of Taipei (Taiwan) Structural designer: Evergreen Consulting Engineering, of Taipei (Taiwan), and Thornton Tomasetti Engineers, of New York City (United States) Damping supplier: Motioneering Inc. Testing laboratory: RWDI Inc. Height: 508 m Interstory height: 4.2 m

FIGURE 8.170 Taipei 101: building overview.

8.2 Mass Damping Systems Case Studies

• • • • •

•

•

Slenderness ratio: 1/10.2 Stories: 101 above grade 1 5 below grade Gross area of the tower: 198,347 m2 Building function: Commercial, office, and hotel space Structural material: Steel super-columns that contain RC up to level 62, involves a total of 95,000 tons of high-strength steel (SM570M), and 23,900 m3 of high-strength concrete (70 MPa) Damping type Tower: Pendulum TMD Pinnacle: Two TMDs Floor plan area: The typical floor plan area is around 2500 m2 (B50 m 3 50 m)

8.2.3.2 Introduction/history Taipei 101 is built in the Hsinyi District of the city, the rapid-growing “Manhattan” of Taipei. The architect, Mr. C.Y. Lee, wished to develop a building form that would reflect the oriental culture. Taipei 101 has eight canted modules bearing on a truncated pyramidal base, square in plan, creating a narrow “waist” at level 26 (Figs. 8.171 and 8.172). Each module with saw tooth corners consists of “8” stories and flares wider at its top, like an opening flower, before the narrow base of the next module starts, forming a setback. The unique profile based on the Chinese lucky number “8,” a homonym for prosperity in Chinese, introduces a new style for skyscrapers. Taipei 101 is so conspicuous not only in its height of 508 m (1667 ft) in 101 stories and unusual building shapes but also in its design to overcome the challenges from potentially extreme earthquakes, frequent and normally severe typhoons and weak subsurface conditions. Consequently, Taipei 101 was elected as one of the Seven Wonders of Engineering by Discovery Channel in 2004 and one of 25 of mankind’s greatest engineering feats by CNN in 2013.

8.2.3.3 Structural system Taipei 101 is mainly composed of a steel structure. To increase structural stiffness as well as strength, the box columns are filled with 69 MPa concrete up to level 62. To prevent soft and weak story at the base of the tower, the steel-braced core is encased in concrete walls from level B5 to level 8. The core tower is relatively compact due to the extensive use of double decked elevators and it is composed of core and perimeter with dual system frames. A square core has four lines of steel diagonal and chevron bracing each way linking 16 core columns. A perimeter framed tube that followed the sloping fac¸ades is used below level 26. Two perimeter super-columns of 2.4 m 3 3 m (7.9 ft 3 9.8 ft) on each face run vertical and change size at upper floors above the 26th floor. The rest perimeter steel moment frames are along each sloping fac¸ade of the building above level 26. Belt trusses at the bottom of each eight-story module are introduced to receive vertical loads from interrupted sloping perimeter columns and deliver

759

760

CHAPTER 8 Case studies of tall buildings

FIGURE 8.171 Taipei 101: typical framing plans.

8.2 Mass Damping Systems Case Studies

FIGURE 8.172 Taipei 101: elevations along interior gridlines.

761

762

CHAPTER 8 Case studies of tall buildings

them to the continuous super-columns. Lateral shear in interrupted perimeter frame also transfers at module setbacks through horizontal steel braces. At selected floors the bracing lines are extended out to engage the perimeter columns through 1-, 2-, and 3-story “outrigger trusses” and form a megaframe. The megastructural system consists of two megaframes in each direction. The refuge mechanical floor on each of the eight modules and other mechanical rooms provide ideal opportunities for numerous well-spaced outriggers. The site is enclosed by a slurry wall 1.2 m (4 ft) thick and an average of 47 m (154 ft) deep to provide sufficient toe embedment below the 21.8
23.5 m (72
77 ft) deep excavation for stability and groundwater cutoff. Within the tower footprint, a continuous RC mat 3
4.7 m (10
15 ft) thick transfers load from column and shear wall loads to a distributed pattern of 380 drilled piles 1.5 m (5 ft) in diameter spaced 4 m (13 ft) on center and socket in bedrock 15
30 m (50
100 ft) deep. The upper portion of Taipei 1010 has eight modules, each with eight stories (Fig. 8.177). Each module flares wider at its top, before the narrow base of the next module. The set of modules bears on a truncated pyramid base, square in plan. A spire rises from the level 91 outdoor observation deck, and is supported on a base of smaller floors flared to echo the sloped walls below. The building frame was to be of structured steel to minimize cost, minimize seismic forces by keeping the building mass low, and benefit from the local steel construction industry. To limit building sway and interstory drift to height 200 in a 50-year storm, stiffness was provided by high-strength concrete. Main building columns up to level 90—16 in the core and two (at upper floors) to four on each face at outrigger lines—are built-up boxes of steel plate. The boxes are filled with 70 MPa concrete up to level 62. The stiffness target was met, having a period of 7 seconds, but occupant comfort was still an issue. Steel framing has low inherent damping, so cyclic wind excitation could build up over time, resulting in accelerations at upper floors reaching uncomfortable levels. The solution was the addition of a TMD (Fig. 8.173). It is interesting to note that for several structural design, local loads are governed by seismic stresses. However, for the overall loading on the structure, i.e., base bending moments, the wind loads govern the design.

8.2.3.3.1 Building fundamental periods The building fundamental vibration modes are shown in Fig. 8.174. The first mode (T1  7.02 seconds) is mainly flexural along Y direction, the second mode (T2  6.96 seconds) is mainly flexural along X direction, and the third one (T3  4.87 seconds) is mainly torsional.

8.2.3.3.2 Damping strategy utilized Wind tunnel tests revealed that, based on building plan dimensions and anticipated wind speeds (3-second gusts of 150 miles per hour at an elevation of 10 m in a 100-year-storm) vortex generation at building corners could occur at a rate matching the tower sway rate or period, causing very large crosswind oscillations.

8.2 Mass Damping Systems Case Studies

Different corner shapes were tested. Double stepped corners brought a dramatic reduction in crosswind excitation and the architect decided to incorporate double steps 2.5 m (8.1 ft) deep at the corners of all eight typical building models (Figs. 8.171 and 8.172). The TMD occupies levels 87
91 as the centerpiece of a public lounge (Fig. 8.173). A sphere of stacked steel with a mass of 725 ton was suspended with eight steel cables looped (criss-crossing) beneath to create a pendulum equivalent to 0.26% of building weight. Adjusting the free cable lengths tunes the sway rate to match the tower (Fig. 8.178). The mass pushes and pulls large dampers as it swings in opposition to the tower.

FIGURE 8.173 Taipei 101: TMD scheme.

763

764

CHAPTER 8 Case studies of tall buildings

FIGURE 8.174 Taipei 101: fundamental frequencies and mode shapes.

Other TMDs address the issue of metal fatigue on the spire. Two compact TMDs were installed to reduce sway on the rooftop spire. Each TMD has a 4.7ton mass sliding horizontally on two sets of orthogonal rollers. Instead of gravity pulling on a pendulum, the mass returns to its central position by spring sets, cables, and pulleys. Dampers link the mass and the building frame.

8.2.3.3.3 Additional damping provided by the damping system Main TMD provides a 3.5% additional damping. Spire TMD 1 has an additional damping of 3.0%, while TMD 2 provides an additional damping of 1.6%.

8.2.3.3.4 Building cost versus damping cost A passive damping system was viewed as a cost-effective means of dealing with the high accelerations.

8.2.3.3.5 Building code Several building codes were utilized and some of them were considered as reference documents. Among all, the most important are: • • • •

Building Code Requirements for Structural Concrete (1997) by Ministry of the Interior, Taiwan Building Code Requirements for Structural Concrete (ACI 318-95 (ACI, 1995)) by American Concrete Institute Foundation Design Specifications for Buildings (1998) by Ministry of the Interior, Taiwan Seismic Design Specifications and Commentary for Buildings (1997) by Ministry of the Interior, Taiwan

8.2 Mass Damping Systems Case Studies

• •

Building Design Code by Ministry of the Interior, Taiwan LRFD Specification for Structural Steel Buildings (1993) by Ministry of the Interior, Taiwan

8.2.3.3.6 Peer-reviewed project The structural peer-reviewed process was carried out by Thornton Tomasetti Engineers (United States).

8.2.3.3.7 Design forces During the strongest wind storm expected to occur in half of a year, according to the Taipei local meteorological records, the building TMD will reduce the peak acceleration of the top occupied floor from 7.9 to 5.0 mg (where 1 mg is 1/1000 of Earth’s standard gravity), to meet the Taiwanese criteria of 5.0 cm/s2 (5.1 mg).

8.2.3.4 Damping overview Located in an adverse construction environment, with significant seismic activity and constant typhoons in season, the Taipei 101 structure required a considerable engineering effort to ensure life safety and comfort. Results of RWDI’s wind engineering studies indicated that accelerations of the building’s upper floors would be 30%
40% higher than desired for this office building. Motioneering was asked to provide the complete damping system for this design/build project, and it spent several months performing the detailed engineering of the TMD, which included determining the behavior of this passive system during extreme seismic events. In addition to designing and supplying the TMD for the tower itself, Motioneering is designing a TMD for the 60 m (197 ft) spire of the same project. As with many spires, fatigue is the prevailing design issue. By increasing the level of energy dissipation in the spire by a significant margin, it is expected that the fatigue life of the spire will be brought into line with the design life of the building itself.

8.2.3.4.1 Damping strategy While TMDs based on swaying pendulums are widely used, they require very tall spaces for mounting and operation, and additional structure for support of their great weight. The simple-pendulum TMD designed to control wind-induced motions is the 660-ton ball-shaped TMD (Fig. 8.175) installed at the top of the Taipei 101 building in Taipei, Taiwan (Haskett et al., 2003; Joseph et al., 2006). A pendulum that spans five floors of the structure. Eight velocity-squared VDDs, designed and manufactured by FIP Industriale S.p.A., Italy, are attached to the TMD mass (Fig. 8.180). The installed length of the VDDs is 3.4 m and their external diameter is 330 mm. The force coefficient of the VDDs is 1000 kN/(m/s)2 and their estimated maximum continuous power dissipation is 25 kW. This TMD is also the first one designed and constructed as a key architectural and visual element in the

765

766

CHAPTER 8 Case studies of tall buildings

FIGURE 8.175 Taipei 101: ball-shaped TMD. Courtesy of CY Shen, T.H. Tsai and Associates, Consulting Engineers.

building. The architect was able to incorporate this vibration absorber into the architectural scheme of the uppermost occupied floors despite the significant amplitude requirements of the TMD under extreme wind and seismic loading scenarios. From the restaurant and bar, through the center of which the TMD penetrates, patrons will be able to see the golden steel ball slightly swinging, many days of the year, under light winds (Fig. 8.180). The TMD is designed to reduce 6-month return period peak accelerations to approximately 5 mg (1 mg is 1/1000 of gravity acceleration), as required by Taiwanese criteria for occupant comfort (Haskett et al., 2003). Since tuning and commissioning of the TMD, several typhoons and earthquakes have hit the Taipei 101 site. The observed performance of both the damper and the tower has been reported to be in-line with analytical predictions and design objectives. For the pinnacle the first approach was to keep the structure as light as possible, in order to minimize demands due to seismic responses for both the pinnacle and the tower structure. With a light pinnacle, the wind-induced vibration and the resultant fatigue from many cycles of this vibration became the dominant design issue. The traditional approach of increasing the mass of the pinnacle would have caused serious implications for the overall tower design with respect to seismic loading. The owner and architect also considered changing the shape to reduce wind effects. Due to the overall structural system for the building, several modes of vibration also included motion of the pinnacle. Three modes (six if counting the perpendicular direction) were found to be affected by vortex-induced vibration (VIO); however, only two were found to be significant relating to fatigue damage. To reduce fatigue damage in the pinnacle, a system of two TMDs was designed and installed by Motioneering. The TMDs was tuned to provide the most benefit to the structure, and they can obtain a significant amplitude reduction in modes 10 and 12. Each TMD weighs 4500 kg (9900 lb) and they are located near the tip of the pinnacle.

8.2 Mass Damping Systems Case Studies

8.2.3.4.2 Damping type At the time of its installation, the Taiwan’s Taipei 101 TMD was the largest pendulum-style TMD ever implemented to control wind-induced motions of a skyscraper. The 660-ton ball-shaped TMD is installed between the 86th and 92nd floors near the top of the tower (Haskett et al., 2003; Joseph et al., 2006). Eight velocity-squared VDDs, designed and manufactured by FIP Industriale S.p.A., Italy, are attached to the TMD mass. The installed length of the VDDs is 3.4 m and their external diameter is 330 mm. The force coefficient of the VDDs is 1000 kN/(m/s)2 and their estimated maximum continuous power dissipation is 25 kW. This TMD is also the first one designed and constructed as a key architectural and visual element in the building. The architect was able to incorporate this vibration absorber into the architectural scheme of the uppermost occupied floors despite the significant amplitude requirements of the TMD under extreme wind and seismic loading scenarios. From the restaurant and bar, through the center of which the TMD penetrates, patrons will be able to see the golden steel ball slightly swinging, many days of the year, under light winds. The TMD is designed to reduce 6-month return period peak accelerations to approximately 5 mg (1 mg is 1/1000 of gravity acceleration), as required by Taiwanese criteria for occupant comfort (Haskett et al., 2003). Pinnacle-tuned mass dampers. Due to the strong wind climate in Taipei, the slenderness of the Pinnacle, and the structural discontinuities at the top of the structure, as it steps down from the full building width to the pinnacle diameter, there exist a number of vibrational modes which cause VIO at common wind speeds. The principal difficulty with VIO in this instance is not that of design wind load, but rather the rapid accumulation of fatigue cycles. VIO has been identified to occur at several frequencies, as follows: 0.656, 0.860, and 1.082 Hz. All of the associated mode shapes involve a simple bending translation of the Pinnacle, with no inflection points. It is the behavior of the levels below the spire base, through to the wide roof and below, that differ in each mode. Also, because of the symmetry of the structure, perpendicular mode shapes are paired at nearidentical frequencies; the pairs of modes 7 and 8, 10 and 11, and 12 and 13 will be hereafter referred to by only the first mode in each pairing. A TMD is very effective in reducing structural response due to narrow banded excitation, as is the case with VIO. On the other hand, the more problematic frequencies (0.86 and 1.08 Hz, corresponding to modes 10 and 12, respectively) are too far apart to be effectively controlled by a single TMD. Therefore, two separate TMDs have been designed, as shown in Fig. 8.176. Each TMD is built to target a different structural frequency and controls the amplitude in both principal perpendicular directions. Fig. 8.177 gives a graphical representation of bins of number of cycles, and the associated base bending moment, induced by VIO in 100 years of wind. This figure also shows the effect of adding the TMDs to the Pinnacle. Note that the number of cycles has not been reduced by the TMDs, but the magnitude of the bending moments has been.

767

768

CHAPTER 8 Case studies of tall buildings

FIGURE 8.176 Taipei 101: dual TMD pinnacle tower.

Given the pinnacle size, height, and wind climate, the quantity of wind energy that is accepted by the structure is very large. In order to achieve the bending moment reductions shown, for the duration of a VIO event that may last for several hours, all of this power must be dissipated by the VDDs as heat. This challenge was met, in the very tight confines shown in Fig. 8.178, by using external fluid circulation in the hydraulic cylinders (VDDs) and a forced air/oil heat exchanger.

8.2.3.4.3 Structural and damping design Code and guidelines. Several building codes were utilized and some of them were considered as reference documents. Among all, the most important are: • •

Building Code Requirements for Structural Concrete (1997) by Ministry of the Interior, Taiwan Building Code Requirements for Structural Concrete (ACI 318-95 (ACI, 1995)) by American Concrete Institute

8.2 Mass Damping Systems Case Studies

FIGURE 8.177 Taipei 101: base bending moment induced by vortex-induced oscillation in 100 years of wind.

FIGURE 8.178 Taipei 101: pinnacle VDD cooling system.

769

770

CHAPTER 8 Case studies of tall buildings

•

Foundation Design Specifications for Buildings (1998) by Ministry of the Interior, Taiwan • Seismic Design Specifications and Commentary for Buildings (1997) by Ministry of the Interior, Taiwan • Building Design Code by Ministry of the Interior, Taiwan • LRFD Specification for Structural Steel Buildings (1993) by Ministry of the Interior, Taiwan Analysis modeling and software. No additional information is available. Design principles. No additional information is available. Design phase considerations. Being a completely passive device means that the building TMD is also in motion during substantially stronger wind events, e.g., 100 years. The design of the TMD must be economically justifiable with regard to possible damage to component parts and the surrounding structure. At such wind levels, the most sensitive devices in this assembly are the VDDs. These VDDs must be able to dissipate, as heat, enough of the energy that they are removing from the structure to avoid overheating and subsequent failure. This is achieved without the use of supplemental liquid cooling by a heat-resistant VDD design (e.g., high-temperature seals and a working fluid which is thermally stable). Extensive testing by the chosen supplier has demonstrated such a level of capability, far beyond the norm in the hydraulic damper industry. The most difficult TMD design aspect was that of handling the seismicity of the region. Due to the differing nature of the building and pinnacle devices, from their size, to their proximity to occupants, completely different seismic design approaches have been taken. The structural engineer provided assistance by sharing an appropriate number of site-specific seismic ground acceleration time histories. These records are utilized to determine the elastic building responses in events up to 100-year return periods. For design events of longer return period (e.g., 1000 years and/or 2500 years), due to the nonlinear (elastoplastic deformation) structural response, time-domain position responses of the floors nearby the TMD were used to evaluate the nuances of building
TMD interaction. The timedomain simulations for these extreme design periods were performed by the structural engineering using DRAIN2D. A rather small conservatism is realized by assuming that the TMD does not reduce the response levels during these design events. Building-tuned mass damper In strong seismic events, for example, those with return periods up to 100 years, the designed response characteristics of the building TMD are quite tame. The steel ring and lower set of eight dampers, shown in Fig. 8.179 on the floor at elevation 374 m, is a secondary system (named a “snubber ring”) designed to engage the TMD only at relative amplitudes which exceed 1 m. While the primary structural response remains linear in seismic events of this strength, the recruitment of this secondary set of VDDs becomes highly nonlinear. To address this computational challenge, a complete kinematic model was assembled to determine the behaviors of TMD/snubber ring collisions, and the sublinear force
velocity damping profiles of the connected secondary VDDs.

8.2 Mass Damping Systems Case Studies

FIGURE 8.179 Taipei 101: main TMD snubber ring.

These lower eight VDDs are drawn from the rail-freight industry, where lowspeed collisions between heavy rail cars are an everyday occurrence. The sublinear force
velocity profile means that upon contact with the snubber ring, at higher stroke velocities, the desired damping force is exerted immediately. As the velocities of the bodies begin to approach each other (a relative velocity approaching zero due to the exerted viscous forces), this preferred level of viscous damping force is maintained with very little change. Substantially more difficulty was encountered when contending with the seismic events of 2500 years’ strength. An interesting observation can be made about the nature of the response of a TMD, tuned only to the first mode of such a slender structure; as many building modes are excited simultaneously, the TMD is seen to respond very little. In essence, the TMD remains almost still in an inertial reference frame, while the building oscillates wildly all around it. In this manner, and due to the phenomenal quantity of energy present in the building during such a strong earthquake, the front-most issue is one of amplitude control (to prevent collision damage), instead of energy dissipation. It was with this in mind that the strength of the snubber system was designed. One difficulty in handling the large viscous forces induced by the snubber ring is the moment it creates on the TMD mass. When the snubber pin engages the snubber ring, well below the TMD, high cable loads are induced on one side of the TMD to resist the overturning of the mass. Conceptually, this occurs as the cables resist the “diving” of the TMD over the snubber ring. The outcome of these types of simulations was incorporated into the design of the supporting wire rope cables and other supporting structural components. Pinnacle-tuned mass dampers In the DE the tip of the Taipei 101 pinnacle is expected to whip back and forth with peak accelerations near 14 g. Under these circumstances, there is no reasonable role that a TMD might be designed to accomplish in terms of vibration reduction. The mass ratio of the pinnacle TMDs, useful for VIO suppression, and nearly a practical maximum given the dimensions

771

772

CHAPTER 8 Case studies of tall buildings

of the pinnacle structure, is completely inadequate to the task of moderating the structural response in strong-to-extreme seismic events. The usual means of handling such excitation do not work in this environment. Enough clearance cannot be given to the TMDs’ translations to prevent extreme collisions with the interior of the pinnacle. Any effort to avoid/restrict large relative TMD motion by increasing the VDD force/velocity coefficient will have a deleterious effect on the performance of the TMDs against VIO. It was at last decided to passively “lock out” the pinnacle TMDs with robust secondary mechanisms. In this manner, the TMDs will travel as inert mass locked to the pinnacle structural system, and avoid any damage that would occur from strong internal collisions—in effect, they will “ride out” the earthquake. This operation happens automatically whenever the TMDs exceed a nominal clearance with sufficient kinetic energy remaining to compress an integrated system of rubber bumpers. “Unlocking,” or freeing this mechanically lockedout state can be accomplished by a single person upon cessation of the seismic event, with the use of further integrated hardware. The ground motion sufficient to cause the pinnacle TMDs to lock out is a peak ground acceleration of approximately 30 cm/ s2. By this estimate, the manual unlocking of the pinnacle TMDs may be required approximately once per year. Most reasonable, a sensor device can be installed to detect this condition and signal the need for an unlocking operation. Designing a passive damping system for both wind and seismic excitation forces the engineer to investigate the problem from two different viewpoints. Generally, the magnitude of forces in wind events tends to be moderate for which designing the system components is readily achievable. However, during extended design wind events, such as typhoons, the amount of energy to dissipated is large and accommodating this within the design can be difficult. Seismic design considerations are generally the opposite, where the total energy to be dissipated is generally more than that for wind, whereas dealing with extremes in force and displacement provides significant challenges for the damping system designer. Taipei 101, where passive TMDs are being implemented to reduce the effects of wind-induced motion on the occupants as well as limiting fatigue damage, is an excellent example of how these two different design considerations are dealt with. By collaborating with both the architect and structural engineer these seemingly conflicting design requirements are effectively integrated—a physical portrayal of the Yin and Yang.

8.2.3.4.4 Architectural integration strategy The view of many owners is that the presence of a special damping device in the building is not something that they necessarily want widely broadcast. In most cases it is tucked away out of view. However, the architects, C.Y. Lee and Associates, and owners of the 101-story Taipei Financial Centre, have taken the route of making the RWDI designed TMD a feature of the building. A special space has been allocated for it near the top of the building so that people will be

8.2 Mass Damping Systems Case Studies

able to walk around it and view it from a variety of angles. It will be brightly colored and special lighting effects are planned. The architect wondered whether the mass had to be of a particular shape, but the design team responded that it could in fact take any shape at all, the only caveat being that a mass of unusual shape might cost more. In the end, a unique concept was proposed: turn the 730-ton (660 mg) TMD into an architectural component with a sculptural dimension: a sphere suspended by flexible steel cables that would be surrounded by three levels of restaurants, bars, and observation decks.

8.2.3.4.5 MEP integration strategy No additional information is available.

8.2.3.4.6 Elevator and other mechanical devices integration strategy No additional information is available.

8.2.3.4.7 Experimental tests RWDI performed wind tunnel testing for the project. A variety of tests were performed to determine the effect of the local wind environment on the pedestrians, for the design of the cladding system and for the structural design of the tower itself. The test results were integrated into the overall scheme of Taipei 101, including a number of minor shape changes to optimize the wind loads with the proposed structural scheme.

8.2.3.4.8 Monitoring system The monitoring system installed in the Taipei 101 tower includes two major parts: the data acquisition and storage platform, sensor subsystem. Sensor subsystems include 30 wired accelerometers which are located at the
5th, 1st, 36th, 60th, 89th, and 101th floors. In order to measure the rotational accelerations of the tower, several accelerators were placed at two locations on a floor along the two major orthogonal axes of this building. Acceleration responses from the tall building were continuously acquired and digitized at 200 Hz by a high-resolution digital data logger. Through the analysis of the data measured from the sensors, structural acceleration responses and dynamic characteristics of the building can be obtained.

8.2.3.5 Comparison of different design strategies 8.2.3.5.1 Structural options considered No additional information is available.

8.2.3.5.2 Damping solution considered No additional information is available.

773

774

CHAPTER 8 Case studies of tall buildings

8.2.3.5.3 Cost
benefit analysis No additional information is available.

8.2.3.6 Lesson learned and recommendations 8.2.3.6.1 Difficulties in the design From a main TMD point of view, the most challenging part of the design was addressing the seismic requirements. Very large amplitudes had to be contained to prevent the TMD from contacting the tower’s super-columns. The pinnacle was also a seismic challenge. With predicted accelerations of 14 g, the loads were tremendous. Also the plan space of 2 m 3 2 m which included the spire structure was very limited. The additional requirement of having maintenance access through the TMDs further limited the space.

8.2.3.6.2 Design innovative solutions No additional information is available.

8.2.3.6.3 Possible improvements No additional information is available.

8.2.4 ONE RINCON HILL (SOUTH TOWER) SAN FRANCISCO, CALIFORNIA, UNITED STATES 8.2.4.1 Project data The major building data (Fig. 8.180) are summarized as follows: • • • • • • • • • • • • •

Year of completion: 2008 Developer: Urban West Associates (United States) Contractor: Bovis Lend Lease (United States) Owner: Urban West Associates (United States) Architectural designer: Solomon Cordwell Buenz United States) and Korth Sunseri Hagey Architects (United States) Structural designer: Magnusson Klemencic Associates (United States) Damping supplier: Bovis Lend Lease (United States) Height: 188 m from street level Interstory height: 2.9 m Stories: highest occupied level 60 with three additional levels of mechanical/ damper space. Ground level varies from level 2 to level 5 Gross area of the tower: 71,063 m2 Building function: Residential Structural material: Concrete classes: 41.4 or 8000 55.2 MPa core walls and column and 37.9 MPa typical slab Rebar: A615 GR60 rebar Prestress: GR270 7-wire prestressing strand Structural steel: 5300 kN capacity BRBs

8.2 Mass Damping Systems Case Studies

FIGURE 8.180 One Rincon Hill: building overview.

•

•

Floor plan area: Typical tower 910 m2 Typical podium: 3530 m2 Damping system: Intrinsic damping: 1.5% Damping type: tuned liquid damper (TLD); 2 tanks, 9.14 m 3 7.62 m 3 1.52 m, 208,652 kg Damper mass to modal mass ratio: 0.56% Added damping: 1.5% Intrinsic 10-year acceleration response: 26 mg Damped 10-year acceleration response: 18 mg

775

776

CHAPTER 8 Case studies of tall buildings

8.2.4.2 Introduction/history One Rincon Hill is a two-tower residential development located in San Francisco, California, United States, on the top of Rincon Hill at the west end of the San Francisco
Oakland Bay Bridge. The south tower, the taller of the two at about 188 m, was the first in an effort to bring high-rise living options to the Downtown San Francisco area. Due to the high seismic risk, BRBs connect the central core shear walls to outrigger columns (Fig. 8.181). In addition to seismic risk, occupant comfort issues were expected under moderate wind events according to the wind tunnel study performed by the Boundary Layer Wind Tunnel Laboratory (BLWTL) at the University of Western Ontario (UWO). These issues are mitigated through the use of a TLD, which reduces accelerations to acceptable levels.

8.2.4.3 Structural system The main characteristics of the building structural system are the following: • • • •

• •

3.66-m thick RC mat foundation 0.25-m thick mild-RC flat slabs spanning 8.53
9.14 m below grade 0.20-m thick post-tensioned concrete flat slabs spanning 8.53
9.14 m above grade Ductile RC core wall system (approximately 17.07 m 3 10.67 m, roughly 10:1 and 15:1 aspect ratios) with supplementary outrigger bracing (reducing aspect ratio from 15:1 to about 8:1) located at levels 28
32 and 51
55 (Fig. 8.182) 0.76 m 3 1.68 m columns supporting BRB outriggers (Fig. 8.183) 2 3 87,064 L capacity TLD tanks located on level 62

8.2.4.3.1 Building fundamental periods The building fundamental periods were both estimated (through numerical analysis, Fig. 8.184) and measured. For the first three modes the values are the following: • •

Estimated: 5.37 seconds (first mode), 4.66 seconds (second mode), and 2.04 seconds (third mode) Measured: 3.86 seconds (first mode), 1.38 seconds (second mode), and 0.90 seconds (third mode)

FIGURE 8.181 One Rincon Hill: (A) south tower typical floor plan, (B) floor plan at BRB, and (C) floor plan at TLD level.

8.2 Mass Damping Systems Case Studies

FIGURE 8.182 One Rincon Hill: south tower lateral system.

8.2.4.3.2 Damping strategy utilized After exploring several options, a tuned liquid sloshing damper (TLSD) was chosen (Fig. 8.187).

8.2.4.3.3 Additional damping provided by the damping system The supplemental system provides an additional 1.5% damping.

8.2.4.3.4 Building cost versus damping cost Cost of damper as a ratio of total construction cost was approximately 1%.

777

FIGURE 8.183 One Rincon Hill: south tower BRB installation.

FIGURE 8.184 One Rincon Hill: fundamental frequencies and mode shapes.

8.2 Mass Damping Systems Case Studies

8.2.4.3.5 Building code The governing building code for this building is the 2001 CBC with the City and County of San Francisco Supplements (CBC, 2001). PBSD was utilized to comply with the alternate design provisions of the building code. Design for the wind was based on wind tunnel study results and since occupant comfort is not addressed in the governing building codes, an acceptable acceleration target of 18 mg at a 10-year recurrence interval was set based on published literature on the subject.

8.2.4.3.6 Peer-reviewed project As a requirement for PBSD, a peer review of the seismic design was performed by a three-person panel consisting of academics and professionals, chaired by Ronald O. Hamburger. The damping system for occupant comfort was not part of the peer review.

8.2.4.4 Damping overview 8.2.4.4.1 Damping strategy Accelerations due to the excitation of a 10-year wind event were determined by wind tunnel testing to be higher than the acceptable level established. To reduce these accelerations, supplemental damping was required; the amount of damping needed was also established by the wind tunnel study. The San Francisco Building Code (SFBC, 2001) requires on-site storage of water for fire protection systems, so this tank was proportioned and located so as to be used as an isolated TLD, resulting in acceptable acceleration levels with minimal architectural impact. The MEP system is unaffected by this strategy, and the water in the tanks is part of the fire protection system. The concept of the TLD is that as the building sways, the force of the water in the tank counteracts the building motion. See Fig. 8.185 for an illustration of how a TLD functions.

FIGURE 8.185 One Rincon Hill: TLD functional scheme.

779

780

CHAPTER 8 Case studies of tall buildings

8.2.4.4.2 Damping type TLSD.

8.2.4.4.3 Structural and damping design As a part of the PBSD, nonlinear time history analyses were performed. These analyses were run using Perform 3D (CSI, 2016). The resulting design takes some exceptions to code-prescribed seismic requirements, particularly with respect to the required secondary lateral system, but was determined to meet or exceed the code intent by a peer-review panel. The building was linearly analyzed for wind using ETABS (CSI, 2016), and the dynamic properties used for evaluating service level wind events were taken from the linear model, as it is expected that the BRBs will behave linearly at this level of excitation. The strength design for wind is code compliant. The benefits of the TLD are not considered in the strength design of the building, though the additional mass is accounted for in the analyses. The design of TLDs was in accordance with the alternate design provision of the code. The level of total damping required, 3%, was determined in the wind tunnel study. The tank was proportioned to satisfy the fire protection requirement and to be able to be tuned to the actual building properties. In addition to an environmental wind analysis as well as wind tunnel studies for strength, occupant comfort, and pedestrian comfort, the UWO BLWTL performed small-scale shake table tests on the TLDs to verify their performance (Fig. 8.186). In situ testing of the as-built building provided the depth of water required in the actual tanks to “tune” the damper to the building.

FIGURE 8.186 One Rincon Hill: shake table test of TLD at UWO.

8.2 Mass Damping Systems Case Studies

Code and guidelines. 2001 California Building Code (CBC, 2001). Analysis modeling and software. The modeling of the building was carried out in ETABS (CSI, 2016) and through in-house spreadsheets. Design principles. The basic principle was to achieve the occupants comfort under service wind load. Design phase considerations. No design phasing was considered.

8.2.4.4.4 Architectural integration strategy TLDs are placed at top-most floors right off the core integrated with architectural and MEP programming.

8.2.4.4.5 MEP integration strategy The water in the TLSD also serves as part of the fire protection system.

8.2.4.4.6 Elevator and other mechanical devices integration strategy There is no interaction between damping device and elevator and other mechanical devices.

8.2.4.4.7 Experimental tests Small-scale shake table tests (Fig. 8.186) were performed to verify the TLDs performance.

8.2.4.4.8 Monitoring system The building is instrumented with a monitoring array of 72 channels of accelerometers (Fig. 8.187) that streams real-time acceleration data from multiple floors starting at level 1 (basement parking level P4) up to roof (Fig. 8.188).

FIGURE 8.187 One Rincon Hill: building accelerometer.

781

782

CHAPTER 8 Case studies of tall buildings

FIGURE 8.188 One Rincon Hill: measured building response to the South Napa earthquake on August 24, 2014.

8.2.4.5 Comparison of different design strategies 8.2.4.5.1 Structural options considered Other structural options such as core with wing walls were considered at early stage of the design but wing walls were replaced with BRB outriggers in response to architectural program.

8.2.4.5.2 Damping solution considered Both TMD and TLD were considered at early stages. TMD was ruled out due to cost implications.

8.2.4.5.3 Cost
Benefit analysis TMD usually occupies smaller areas as opposed to TLD. A cost
benefit analysis was performed comparing the net salable area gained by using TMD. Considering the dual application of water tank for both fire protection system and added damping, TLD was selected as final option.

8.2 Mass Damping Systems Case Studies

Damping cost. Cost of damper as ratio of total construction cost was approximately 1%.

8.2.4.6 Lesson learned and recommendations 8.2.4.6.1 Difficulties in the design Given the high seismicity of San Francisco, additional mass at the top-most levels added extra lateral seismic load on the building.

8.2.4.6.2 Design innovative solutions This project has the first application of BRB as outrigger elements in a tall building.

8.2.4.6.3 Possible improvements No possible improvements could have been considered.

8.2.5 COMCAST CENTER, PHILADELPHIA, PENNSYLVANIA, UNITED STATES 8.2.5.1 Project data The major building data (Fig. 8.189) are summarized as follows: • • • • • • • • • • • • • • •

Year of completion: 2008 Developer/contractor: Liberty/Commerz (United States) Architectural designer: Robert A.M. Stern (United States) Structural designer: Thornton Tomasetti Engineers (United States) Damping supplier: NA Testing laboratory: RWDI Inc. (Canada) Height: 296.7 m Interstory height: 4.50 m Slenderness ratio: 1/5.8 Stories: 57 above grade 1 3 below grade Gross area of the tower: 297,000 m2 Building function: Commercial and offices Structural material: No information available Damping type: Tuned liquid column damper Floor plan area: 2140 and 2600 m2

8.2.5.2 Introduction/history Comcast Center is an office tower that includes restaurant and retail space. It features a double-skin atrium curtain wall with active air control at the 30-m tall main lobby. The tower tapers at the top with two cutouts near the top of the building.

8.2.5.3 Structural system Central concrete core with steel-framed floors.

783

784

CHAPTER 8 Case studies of tall buildings

FIGURE 8.189 Comcast Center: building overview.

8.2 Mass Damping Systems Case Studies

8.2.5.3.1 Building fundamental periods The Comcast Center is expected to oscillate once every 7 seconds when deflected by wind. A 33.5-m/s wind would move the top floors about 46 cm.

8.2.5.3.2 Damping strategy utilized Tuned liquid column dampers (TLCDs) are special types of dampers relying on the motion of a column of liquid in a U-shaped container to counteract the forces acting on the structure. Damping is introduced in the oscillating column through an orifice in the liquid passage. The water will slosh in the opposite direction in the tank’s twin U-shaped chambers. In water tanks like the Comcast Center’s, the energy is dissipated by vertical steel vanes, or louvers, that impede the back-andforth flow of water and cause turbulence in the flow. The TLCD is a 1179-ton device (with 1,135,625 L of water) designed to control the structure’s user comfort and floor-to-floor lateral drift. The design features two chambers for extra efficiency. In high winds, the TLCDs reduce by about one-third the acceleration that a person would feel on the upper floors (Fig. 8.190). To optimize lateral drift serviceability performance, the tower required supplementary damping in its most slender axis only. Some custom features of this uniaxial TLCD include: • • •

TLCD water mass 5 1179 tons (over 1,135,625 L of water) TLCD is almost as wide as it is long—an unusual feature Tank divided into two to avoid off-axis flow (loss of efficiency)

8.2.5.3.3 Additional damping provided by the damping system Both occupant comfort and drift criteria for the building were realized through the use of the world’s largest TLCD.

8.2.5.3.4 Building cost versus damping cost Overall structural cost savings (in the millions of dollars) were realized by using TLCDs to achieve occupant comfort and lateral drift serviceability criteria.

FIGURE 8.190 Comcast Center: TLCD schematic view.

785

786

CHAPTER 8 Case studies of tall buildings

TLCD construction costs were minimized since Motioneering sourced most of the hardware on the owners behalf. Use of the TLCD resulted in reduced overall construction costs, as less structural steel was required. Engineers around the world have begun to choose water for some damping projects in part because it is cheaper than steel, and because pumping water to the top of a building is easy.

8.2.5.3.5 Building code No information is available.

8.2.5.3.6 Peer-reviewed project No information is available.

8.2.5.3.7 Expected performance In high winds, the TLCDs reduce by approximately one-third the acceleration that a person would feel on the upper floors.

8.2.5.4 Damping overview The 307-m tall Comcast Center contains the largest TLCD in the world (at the time of construction).

8.2.5.4.1 Damping strategy To optimize lateral drift serviceability performance, the tower required supplementary damping in its most slender axis only. Therefore, one large TLCD was designed instead of the more typical two perpendicular TLCD configurations.

8.2.5.4.2 Damping type Some custom features of the uniaxial TLCD include: • • •

TLCD water mass 5 1,135,00 L TLCD is almost as wide as it is long—an unusual feature Tank divided into two to avoid off-axis flow (loss of efficiency)

8.2.5.4.3 Structural and damping design In water tanks like the Comcast Center’s, the energy is dissipated by vertical steel vanes, or louvers, that impede the back-and-forth flow of water. Code and guidelines. No information is available. Analysis modeling and software. The wind tunnel data files collected during scale model testing were used as the excitation force in nonlinear numerical simulations, performed by custom RWDI in-house software. Performance and loads were determined resulting from both service level winds, and for ultimate design loads (both wind and seismic). Design principles. No information is available.

8.2 Mass Damping Systems Case Studies

Design phase considerations. In the early design stages of the project, space was allocated at the top of the building for the supplementary damping system.

8.2.5.4.4 Architectural integration strategy No information is available.

8.2.5.4.5 MEP integration strategy No information is available.

8.2.5.4.6 Elevator and other mechanical devices integration strategy No information is available.

8.2.5.4.7 Experimental tests Full-scale excitation of the completed TLCDs was conducted by pressurizing one end of each TLCD to displace the water, and then suddenly releasing this pressure. Upon release, the water oscillates and interacts with the structure, and the entire response is measured by accelerometer and analyzed. Successive tests were conducted for different openness (porosity) settings of the louver dampers, to map the partial/internal damping of the TLCD to louver settings. The optimal setting was then instituted, to match the best wind serviceability response determined in initial studies.

8.2.5.4.8 Monitoring system No information is available.

8.2.5.4.9 Maintenance strategy Periodically (e.g., every couple of years) the water in a single TLCD is to be drained, so that the interior of the TLCD can be inspected. The integrity of the waterproofing membrane is inspected, and the drag-inducing louvers are inspected for corrosion and have protective paint reapplied if necessary. Upon completion, the water in the second TLCD can be pumped over to the first, and then the second TLCD is similarly inspected. This retains most of the damping performance so that moderate winds do not cause sufficient motion for adverse comment, and minimizes the amount of water lost by draining to sewer.

8.2.5.5 Comparison of different design strategies 8.2.5.5.1 Structural options considered No information is available.

8.2.5.5.2 Damping solution considered No information is available.

787

788

CHAPTER 8 Case studies of tall buildings

8.2.5.5.3 Cost
benefit analysis Overall structural cost savings (in the millions of dollars) were realized by using TLCDs to achieve occupant comfort and lateral drift serviceability criteria. Damping cost. Damper budget cost was in the order of $2 million (source DHUY.com).

8.2.5.6 Lesson learned and recommendations 8.2.5.6.1 Difficulties in the design The contractor wished to use precast concrete for the construction of the TLCD, making successful waterproofing more difficult than if it had been cast-in-place.

8.2.5.6.2 Design innovative solutions No information is available.

8.2.5.6.3 Possible improvements No information is available.

8.2.6 HYATT PARK TOWER, CHICAGO, ILLINOIS, UNITED STATES 8.2.6.1 Project data The major building data (Fig. 8.191) are summarized as follows: • • • • • • • • • • • • • •

Year of completion. 1999 Developer/contractor: Hyatt Development Corporation, Park Tower, LLC (United States) Architectural designer: Lucien Lagrange and Associates, HKS Inc. (United States) Structural designer: Chris Stefanos and Associates (United States) Damping supplier: RWDI (Canada)/Taylor Devices Inc./McHugh Construction Testing laboratory: Wind tunnel tests performed at RWDI Inc. (Canada) Height: 239-m top floor height, 250-m top of roof, and 257-m top spire Interstory height: 4.50 m Stories: 67 above grade 1 1 below grade Gross area of the tower: 76,924 m2 Building function: Condominium/hotel Structural material: No information available Damping type: Pendulum TMD Floor plan area: No information available

8.2.6.2 Introduction/history No information is available.

8.2 Mass Damping Systems Case Studies

FIGURE 8.191 Hyatt Park Tower: building overview. © Lucien Lagrange Architects/CTBUH.

789

790

CHAPTER 8 Case studies of tall buildings

8.2.6.3 Structural system RC shear wall core with perimeter tube frame of spandrel beams and columns. High-strength concrete used for primary lateral load resisting system.

8.2.6.3.1 Building fundamental periods The Park Tower has a fundamental period of 4.23 seconds (X direction, first mode), 3.92 seconds (Y direction, second mode), and 1.72 seconds (torsion, third mode).

8.2.6.3.2 Damping strategy utilized The wind tunnel tests showed that with the initially planned structural system the accelerations would be above the desired values for a residential building. The 10-year return period acceleration was predicted to be in the range of 26
30 mg at 2% damping ratio and a target of 15 mg had been set. The higher-than-normal accelerations were primarily due to wind flows off the John Hancock Tower nearby. After initial investigations indicated that the peak acceleration of the upper floors of the Park Tower would have exceeded the maximum target value, a solution was proposed that included incorporating a TMD near the top of the building. It consists of a 300-ton mass block suspended from cables with lengths adjustable up to 8.0 m. This mass represented approximately 2.0% of the building’s generalized mass in the first mode of vibration. The tuned mass is damped using four special hydraulic dampers that are hermetically sealed, frictionless, and are designed to have an unlimited and maintenance-free life for the life of the building. The TMD is equipped with an antiyaw device and a frequency-tuning frame to allow adjustment after the building was completed. Additionally, the TMD is equipped with a pair of hydraulic brakes that are also hermetically sealed. The brakes have control valves that can be closed in a controlled manner to bring the mass to a rest and keep it locked in place as necessary. Separate overtravel buffers are also installed to provide additional energy absorbing capability in the event that the tuned mass exceeds its expected amplitude (e.g., wind events with probability of less than 1-in-50 years).

8.2.6.3.3 Additional damping provided by the damping system The additional damping provided was 3.0%.

8.2.6.3.4 Building cost versus damping cost After extensive investigations into various structural or shape change solutions, the decision was made to add damping. From the developer’s point of view the damping system had to be economical, require little maintenance, and avoid

8.2 Mass Damping Systems Case Studies

compromising valuable floor space. The design selected consisted of a simple pendulum damper mounted under the mansard roof of the tower, an area being used for mechanical equipment. With some minor changes in the geometry of the mansard, there was enough space to accommodate a simple pendulum TMD. The simplicity of the design minimizes the need for maintenance and also kept the cost low. The only components in need of eventual maintenance will be the hydraulic dampers and present-day hydraulic dampers can be manufactured to have very low maintenance. The TMD cost: four hydraulic dampers, two hydraulic brakes, four overtravel buffers were supplied for $369,000. Data on the fabrication and installation cost of the steel box and plates, cables, is not available.

8.2.6.3.5 Building code No information is available.

8.2.6.3.6 Peer-reviewed project No information is available.

8.2.6.3.7 Expected performance Initial lateral analysis: • • •

ASCE
Exposure A T 5 7.26 seconds Drift 5 340 mm 5 h/(682) Acceleration predictions:

•

High-frequency force balance (HFFB) result: 26.1 mg Accepted structural modifications:

• • •

Increase link beam depth Increase perimeter spandrels Increase perimeter columns in tube Modified structure:

• • •

T 5 4.7 seconds Drift 5 246 mm 5 h/940 Accelerations 5 20.7 mg (10-year return period)

No further changes to the structure could be achieved without imposing on the architecture. The solution to high accelerations was to increase damping. In order to quantify the overall improvement realized through the addition of the TMD, we can compare the peak acceleration of the top occupied condominium of the building with respect to the return period both with and without the TMD. Note that the reduction in acceleration is significant (Fig. 8.192). In fact, recall that the criteria suggested by the International Organization for

791

CHAPTER 8 Case studies of tall buildings

3 Total acceleration (% g)

792

2.5 2 1.5 1 0.5 0 0.1

1 Return period (years) ISO criteria*

Without TMD

10

With TMD

* Factored for residential occupancy

FIGURE 8.192 Hyatt Park Tower: TMD acceleration performance.

Standardization (ISO) over a 5-year return period is approximately 18 mg (frequency-dependent ISO criteria in effect at the time). The addition of the TMD resulted in a peak acceleration of less than 15 mg over a 10-year return period. The addition of the TMD resulted in lowering the acceleration at the top condominium from approximately 12 mg to 8 mg for the 1-year return period and approximately 21 mg to less than 15 mg for the 10-year return period. Since this TMD is meant to address acceleration due to wind only, seismic performance was not a concern.

8.2.6.4 Damping overview The Park Tower would not have needed a damping system except for its proximity to the 100-story John Hancock Center. Wind blowing off Lake Michigan swirls around the Hancock building, creating high-speed venture gusts that rock the Park Tower structure. The 300-ton damper, which is 2.4 m wide, 5.5 m long, and 3.4 m high, is made up of 50-mm thick steel plates stacked on top of each other. Features of the TMD are that it has a tuning frame which can be moved up and down and clamped on the cables to allow the natural period of the pendulum to be adjusted. The damper constants can be adjusted. The dampers are nonlinear with a force proportional to velocity squared so as to prevent excessive mass motions during extreme wind events. The mass is connected to an antiyaw device to prevent rotations about a vertical axis. Below the mass there is a bumper ring connected to hydraulic buffers to prevent travel beyond the hydraulic cylinder length. The main hydraulic dampers of the TMD are sloped from their floor

8.2 Mass Damping Systems Case Studies

mountings up to the TMD mass. This was found to be advantageous in shortening the stroke required of the dampers which reduced cost. The building frequencies have been measured using accelerometers mounted in the building and by recording motions caused by ambient winds. The measured sway frequencies were within 10%
20% of computed modal frequencies. The results of the measurements will allow the correct pendulum length to be set. The predicted 10-year acceleration for the building with the TMD in operation is 15 mg. For a building like the Park Tower, with a period of about 5 seconds, this translates into acceleration criteria of approximately 12 mg (1.2% of gravity) for the 1-year return period and 18 mg for the 5-year return period. Factoring the ISO criterion to account for the residential occupancy and extrapolating the 5-year return period to the 10-year return period, a generally accepted criterion of 15 mg (1.5%) of gravity is produced.

8.2.6.4.1 Damping strategy In an effort to reduce the wind-induced moments and accelerations, the structural engineers investigated a number of possible structural solutions. Additional mass was added by thickening all of the floor slabs while additional stiffness was provided by deepening the spandrel beams. All of these structural measures reduced the resonant effects, but they were not sufficient to reduce the accelerations to an acceptable level. Any additional structural modifications would have impacted the architecture of the Park Tower and further work on structural modifications was abandoned in favor of investigating supplementing the structural damping of the building. A TMD was selected since the acceleration only needed to be reduced a limited amount (Table 8.37). The required damping was not for structural load or seismic performance purposes. Conceptual design of TMD: • •

Mass required to reduce accelerations to target Space requirements (vertical and horizontal) necessary for operation of damper Table 8.37 Hyatt Park Tower: TMD Characteristics Characteristic

Value

Mass Peak damping force (per direction) X displacement (500 years) Y displacement (500 years) Resultant displacement (500 years) Length Maximum pendulum length Minimum pendulum length

340 tons 178 kN 0.90 m 0.80 m 0.76 m 7.9 m 5.8 m

793

794

CHAPTER 8 Case studies of tall buildings

• • • • • • • • •

TMD was to be a passive device Low maintenance Detailed design of TMD Viscous damper design Antiyaw device to prevent torsional motion of damper Adjustable pendulum length Adjustable damper force Nonlinear damper force to prevent excessive motion during extreme events Bumper ring to prevent travel beyond cylinder length

8.2.6.4.2 Damping type A number of damping systems were investigated during a concept design phase. A TMD was selected as space was available for such a device in the mechanical levels at the top of the tower, and also the proposed concept was a simple device with no maintenance requirements, other than routine inspection. In reality, structures are rarely linear and it may also be advantageous to use nonlinear characteristics of certain components in designing for more severe wind events. This is the case for the VDDs used for the TMD damping on the Park Tower, where the higher forces produced by the nonlinear VDDs beneficially reduce the peak TMD displacements under the most severe wind events. Fig. 8.193 depicts the final configuration of the TMD that was designed by RWDI. The original illustration in color highlights a number of components. The main support cables are paired at each corner of the rectangular mass block. The structural steel above the mass block (blue and red in the color image), near the top of the image, is the tuning frame that can be lifted or lowered as required to tune the

FIGURE 8.193 Hyatt Park Tower: TMD scheme.

8.2 Mass Damping Systems Case Studies

TMD to the frequency of the building. The red frames depicted in the foreground of the image (Fig. 8.198) are a mechanism called an antiyaw mechanism. This device prevents the TMD mass from rotating (or yawing) as it swings from side to side. Two viscous dampers are also connected to the antiyaw mechanism that will be used as a brake if it becomes necessary to stop the motion of the TMD. The angled cylindrical devices that slope from the mass to the floor are the viscous dampers specified by RWDI and supplied by Taylor Devices. TMDs are typically required to consist of dampers with stringent output requirements in order to optimize the performance. These requirements typically include: • • • • • •

Proper damping function whereby the output force is dependent on velocity Low friction Maintenance-free design High continuous power output for severe input conditions High damping forces to properly control the mass and dissipate energy Tightly controlled performance parameters overall operating conditions

The hydraulic damper design selected for the Park Tower was a patented maintenance-free low friction, hermetically sealed fluid damper designed, and manufactured by Taylor Devices, Inc. For the purposes of this project, damper output force varies with velocity squared. Although not always necessary, a damper that obeys this relationship has the advantage of being able to react favorably to high wind levels. During rare storms when the velocity of the tuned mass is high, the dampers absorb much more energy and can therefore prevent excessive motion of the tuned mass. Although somewhat easier to model, linear systems do not have this advantage and may therefore need to compensate for extreme input levels in some other fashion. The hydraulic dampers were designed to produce forces as high as 44.5 kN each and absorb 1.8 horsepower during the most severe wind conditions.

8.2.6.4.3 Structural and damping design Code and guidelines. No information is available. Analysis modeling and software. To determine structural wind loads and windinduced accelerations, a scaled model of the building as shown in Fig. 8.194 is placed in the wind tunnel and tested for 36 azimuthal wind directions. The immediate surroundings of the building are modeled accurately, while the upwind fetches are modeled to generate the appropriate wind profile. For example, winds approaching the building over the waters of Lake Michigan are modeled as an ASCE profile “C,” while winds approaching the building from the west are modeled as ASCE profile “B.” In the wind tunnel, allowances can be made for additional roughness where an upwind profile might be somewhere between an ASCE profile “B” and “A.” An HFFB test was performed where the bending moments and shears are measured directly from the model to give generalized wind forces. These are then combined with structural

795

796

CHAPTER 8 Case studies of tall buildings

FIGURE 8.194 Hyatt Park Tower: wind tunnel test.

characteristics, such as building mass, mode shapes, and an estimate of the structural damping to determine the full-scale dynamic behavior of the building. Additional wind tunnel testing using an aeroelastic wind tunnel model was performed to obtain more accurate predictions of the building accelerations. Aeroelastic wind tunnel testing allows one to measure the dynamic response directly from the model and will account for the usually beneficial effect of aerodynamic damping as well as a more accurate characterization of the peak responses. Fig. 8.195 shows a particular result of such a wind tunnel test for the Park Tower. The wind-induced bending moment in one of the building’s principal axes, that include the mean value, background, and resonant components are plotted versus 36 azimuthal wind directions. It is interesting to note that there is a range of wind directions from approximately 10 and 40 degrees (as measured from north) where the wind-induced bending moments are much larger than those for the remaining wind directions. Further investigations into these wind directions revealed that upwind buildings were having a significant impact on the wind-induced responses of the building, as these higher bending moments are indicative of higher accelerations as well. Fig. 8.195 also shows the results of the aeroelastic wind tunnel testing, that for most wind directions indicated reduced bending moments and correspondingly lower accelerations. The final predicted accelerations for the Park Tower were in the range of 20
23 mg, which was higher than the desired criterion of 15 mg. Design principles. In order to qualify the power dissipating capability of the damper, an input file was programmed into the test machine that represented the maximum expected power level. The temperature of the damper was then

8.2 Mass Damping Systems Case Studies

FIGURE 8.195 Hyatt Park Tower: wind-induced bending moments.

recorded after a condition of steady-state heat transfer was achieved. In this case, the damper was expected to be capable of dissipating almost 2 horsepower with no external means of cooling allowable. Since the design of the damper accommodates a higher working temperature than conventional dampers, the power input level proved to be well within the capability of the damper. As previously mentioned, the damper design involves a crossover port that ensures equal distribution of heating on the external surfaces of the damper. During power verification testing, the damper exhibited a very consistent and well-balanced temperature rise along its entire length, thereby efficiently dissipating the heat to the surrounding environment. Another critical component of the TMD is the set of eight support cables. These cables not only suspend the 272-ton mass block, but also are of an appropriate length to allow the mass to behave as a natural pendulum. Cables were selected over other possible vertical supports, as a rotational bearing need not be provided at either end of the support when using a cable. As the mass will swing from side to side, the cables will have to flex to allow this movement. RWDI designed special cable housings with appropriately machined surfaces for the cables to bend over as the mass swings to either side. These cable housings also ensured that the location where the bending stresses are induced into the cables differ from the end connection of the cable. Given the relatively large diameter of the support cables of 63.5 mm, spelter fittings were used for the end connections of the cables. In addition, an ultraflexible cable design was used to assist with reducing bending stresses in the cables which could over the long

797

CHAPTER 8 Case studies of tall buildings

term cause fatigue damage to the cables. To verify the factors of safety for the design of the cables, a cable destruction test was performed on a ninth cable. This test demonstrated that the design factor of safety of 3 for each cable was exceeded by 40%. In reality the design factor of safety for the support system is closer to 6 as RWDI designed the cable support system in such a way that only one cable is required at each corner. This way, if one cable were to fail, there would be a backup cable. The 63.5-mm-diameter cable failed under a load of over 282,588 kg (272 tons). The results of the testing provided an excellent level of comfort for the cable system that supports the 272-ton mass block. Design phase considerations. A number of damping systems were investigated during a concept design phase. A TMD was selected as space was available for such a device in the mechanical levels at the top of the tower, and also the proposed concept was a simple device with no maintenance requirements, other than routine inspection. Den Hartog (1956) and others have developed the background theory for a TMD to produce simple expressions predicting the behavior of a structure with a supplemental moving mass attached (see Chapter 4 and Appendix A). Using these simple expressions, it is possible to predict the effective damping of a structure with a TMD installed. This effective damping is a combination of the original structural damping and the damping provided by the TMD. These predictions can be made as a function of the size of the added mass as a ratio of the generalized mass of the building and the frequency or tuning ratio. The effective damping that can be provided using a TMD for a number of mass ratios is given in Fig. 8.196. 7 Effective damping (%)

798

6 5 4 3 2 1 0 0.7

0.8

0.9 1 1.1 Frequency or tuning ratio Mass ratio 1.00%

2.00%

5.00%

Effectiveness of a TMD

FIGURE 8.196 Hyatt Park Tower: wind-induced bending moments.

1.2

1.3

8.2 Mass Damping Systems Case Studies

It should be noted that these expressions and relationships are for a linear system only. In reality, structures are rarely linear and it may also be advantageous to use nonlinear characteristics of certain components in designing for more severe wind events. This is the case for the VDDs used for the TMD damping on the Park Tower, where the higher forces produced by the nonlinear VDDs beneficially reduce the peak TMD displacements under the most severe wind events.

8.2.6.4.4 Architectural integration strategy No information is available.

8.2.6.4.5 MEP integration strategy No information is available.

8.2.6.4.6 Elevator and other mechanical devices integration strategy No information is available.

8.2.6.4.7 Experimental tests Hydraulic dampers were component level tested to verify force versus velocity function throughout the velocity range, the continuous power capability, and the friction. The frequency of the finished structure was measured and the frequency of the TMD was adjusted using the tuning frame accordingly. An in-depth inspection and repeat of commissioning pull tests were again conducted in 2010. The building natural frequencies were found to have decreased only slightly over a decade, approximately 3.2% in the EW axis, and 1.3% in the NS axis. No adjustment to the tuned frequency of the TMD was required, because it had originally been tuned slightly low for the as-built frequencies of the new building, and was now in nearly perfect alignment with the newly measured in-service frequencies (as intended), with a low amplitude structural damping of almost 0.8%.

8.2.6.4.8 Monitoring system No information is available.

8.2.6.4.9 Maintenance strategy As stated earlier, the hydraulic dampers were designed to be completely maintenance free for the life of the building. Occasional routine inspection of the TMD itself has been performed on an ongoing basis with no problems noted.

8.2.6.5 Comparison of different design strategies After extensive investigations into various structural or shape change solutions, the decision was made to add damping. From the developer’s point of view the damping system had to be economical, require little maintenance, and avoid compromising valuable floor space. The design selected consisted of a simple pendulum damper mounted under the mansard roof of the tower, an area being

799

800

CHAPTER 8 Case studies of tall buildings

used for mechanical equipment. With some minor changes in the geometry of the mansard, there was enough space to accommodate a simple pendulum TMD. It consists of a 272-ton mass block slung from cables with lengths adjustable up to 10.5 m. This mass represented approximately 1.4% of the building’s generalized mass in the first mode of vibration. The simplicity of the design minimizes the need for maintenance and also kept the cost low.

8.2.6.5.1 Structural options considered Damping can be added to a structure through various means. VE dampers or linear hydraulic dampers have been incorporated in some structures in locations where relative motions exist, such as at beam
column connections. Using the relative motion, energy can be absorbed as heat created by the relative motion in the damper itself. For the Park Tower, due to its construction in concrete and its relatively high stiffness, the use of this type of damper was not expected to be as effective as other methods. The use of added mass dampers at the top of the tower was seen as a more effective route. Sloshing liquid dampers are one option, but due to the relatively low density of water and the reduced effectiveness of sloshing type action, this approach would require an excessive volume of water. To reduce the size of the overall package, a more dense material such as steel is required; hence the use of a passive TMD was investigated. As there was sufficient clear height available in the roof volume, a pendulum-type passive damper was selected. Several semiactive TMDs exist in North America, but operational and reliability complications of a semiactive system would not be cost-effective for this application.

8.2.6.5.2 Damping solution considered No information is available.

8.2.6.5.3 Cost
benefit analysis Damping cost. The cost of the hydraulic dampers, hydraulic brakes, and the overtravel buffers was approximately $369,000.

8.2.6.6 Lesson learned and recommendations 8.2.6.6.1 Difficulties in the design No information is available.

8.2.6.6.2 Design innovative solutions Anticipating that the as-built frequencies of the tower would be dissimilar to those predicted by computer model, a frame for adjusting the TMD pendulum frequency was designed so that it could be properly tuned to the as-built structural frequencies.

8.2 Mass Damping Systems Case Studies

8.2.6.6.3 Possible improvements No information is available.

8.2.7 HIGHCLIFF APARTMENTS, HONG KONG 8.2.7.1 Project data The major building data (Fig. 8.197) are summarized as follows: • •

• • • • • • • • •

• •

Year of completion: 2003 Developer/contractor: Central Development Limited (Hong Kong) Hip Hing Construction Company (Hong Kong) Architectural designer: Dennis Lau & Ng Chun Man Architects and Engineers (Hong Kong) Structural designer: Magnusson Klemencic Associates (United States) Damping supplier: Tank constructed by the main contractor Height: 252.3 Interstory height: 3.15 m Stories: 73 above grade 1 7 below grade Gross area of the tower: 34,931 m2 Building function: Residential Structural material: Concrete classes: Columns and shear walls: 65 MPa; slab: 45 MPa; rebar: BS 4449 deformed bars, type T, grade 460 rebar Floor plan area: Two offset intersecting ellipses at a total of about 18 m deep 3 48 m long Damping system Intrinsic damping: 1.5% Damping type: TLD; 4 tanks, 9.67 m 3 3.6 mx1.2 m, 34,812 kg of water Damper mass to modal mass ratio: 0.34% Added damping: 1.6% Intrinsic 10-year acceleration response: 24 mg Damped 10-year acceleration response: 15 mg

8.2.7.2 Introduction/history Highcliff Apartments is a residential tower located on a steep hillside in Hong Kong. This hillside is considered one of the windiest places on earth. Additionally, the tower is quite slender, with an aspect ratio of just 12 to 1, making it currently one of the most slender buildings in the world. To provide strength and improve economy, all vertical elements of the building are engaged by the lateral system. Additionally, TLDs are located at the top of the building to reduce accelerations under the service level wind event to maintain occupant comfort.

801

802

CHAPTER 8 Case studies of tall buildings

FIGURE 8.197 Hyatt Park Tower: building overview.

8.2 Mass Damping Systems Case Studies

FIGURE 8.198 Hyatt Park Tower: typical tower floor.

8.2.7.3 Structural system The tower is a reinforced high-strength concrete structure that efficiently utilizes all structural elements in the tower both as its gravity and lateral force resisting systems. The tower floor plate is composed in plan view of two offset intersecting ellipses (Fig. 8.198), a total of about 18 m deep 3 48 m long. At the center of this intersection and toward the back-of-house are RC core walls that are joined by a concrete dividing wall to a large column on the east face. This wall has doorways to allow flexibility of access between units, but the doorways are offset from floor to floor to allow the wall to retain its needed stiffness. Within the apartment units, dividing walls are also used to increase the stiffness of the slender direction. A continuous RC beam, approximately 1 m deep 3 0.300 m wide, runs around the perimeter of the floor. This beam not only supports the concrete flat-slab floor framing but also effectively engages all lateral resisting elements around the perimeter to resist the torsional affects due to wind. At each of the two refuge floors located at about the third points of the tower, continuous walls with openings replace the perimeter beams. These “belts” provide further stiffness to the tower. Similarly, perimeter walls at the roof form “hat” that also helps to resist the wind forces. Within the mechanical penthouse of the roof, four sloshing dampers mitigate acceleration due to frequent winds (Fig. 8.199).

8.2.7.3.1 Building fundamental periods The building fundamental periods were both estimated (from numerical analysis) and measured, as follows (Fig. 8.200): • •

Estimated: 6.29 seconds (first mode), 4.30 seconds (second mode), and 2.38 seconds (third mode) Measured: 6 seconds (first mode) and 3.8 seconds (second mode)

8.2.7.3.2 Damping strategy utilized After exploring several options, a TLSD was chosen (Fig. 8.201).

803

804

CHAPTER 8 Case studies of tall buildings

FIGURE 8.199 Hyatt Park Tower: floor plan at mechanical penthouse showing damper tanks and “hat” perimeter walls.

FIGURE 8.200 Hyatt Park Tower: fundamental frequencies and mode shapes.

8.2.7.3.3 Additional damping provided by the damping system The supplemental damping system adds an additional 1.6% damping.

8.2.7.3.4 Building cost versus damping cost Cost of damper as a ratio of total construction cost was approximately 1%.

8.2.7.3.5 Building code No reference code is available for sloshing damper.

8.2 Mass Damping Systems Case Studies

FIGURE 8.201 Hyatt Park Tower: TLD schematic illustration.

8.2.7.3.6 Peer-reviewed project The project was not peer reviewed.

8.2.7.4 Damping overview 8.2.7.4.1 Damping strategy Accelerations due to the excitation of a 10-year wind event were determined by wind tunnel testing to be higher than the acceptable level established. To reduce these accelerations, supplemental damping was required; the amount of damping needed was also established by the wind tunnel study. The Hong Kong Building Code requires on-site storage of water for fire protection systems, so this tank was proportioned and located to also be used as an isolated TLD, resulting in acceptable acceleration levels with minimal architectural impact. The MEP system is unaffected by this strategy, and the water in the tanks is part of the fire protection system. The concept of the TLD is that as the building sways, the force of the water in the tank counteracts the building motion. See Fig. 8.206 for an illustration of how a TLD functions.

8.2.7.4.2 Damping type TLSD.

8.2.7.4.3 Structural and damping design Code and guidelines. There is no code for the design of sloshing dampers. The level of total damping required, 3%, was determined in the wind tunnel study. The tank was proportioned to satisfy the fire protection requirement and to be able to be tuned to the actual building properties. In addition to an environmental wind analysis as well as wind tunnel studies for strength, occupant comfort, and pedestrian comfort, the UWO BLWTL performed small-scale shake table tests on the sloshing dampers to verify their performance. In situ testing of the as-built building provided the

805

806

CHAPTER 8 Case studies of tall buildings

FIGURE 8.202 Hyatt Park Tower: TLD internal view of dissipating screens.

depth of water required in the actual tanks to “tune” the damper to the building. Fig. 8.202 shows a photographs of the inside of one of the tanks. Analysis modeling and software. The building was linearly analyzed using ETABS (CSI, 2016), and the dynamic properties used for evaluating service level wind events were taken from this model. The benefits of the sloshing damper are not considered in the strength design of the building, though the additional mass is accounted for in the analyses. Design principles. The basic principle was to achieve the occupants comfort under service level wind load. Design phase considerations. No design phasing was considered.

8.2.7.4.4 Architectural integration strategy TLDs are placed at topmost floors right off the core integrated with architectural and MEP programming.

8.2.7.4.5 MEP integration strategy The water in the TLSD also serves as part of the fire protection system.

8.2.7.4.6 Elevator and other mechanical devices integration strategy There is no interaction between damping device and elevator and other mechanical devices.

8.2.7.4.7 Experimental tests No experimental test was performed.

8.2.7.4.8 Monitoring system To ensure damper performance, a comprehensive monitoring system was installed by Magnusson Klemencic Associates. Sensors located near the roof measure the building motion (Fig. 8.203) and wave action in the tanks (Fig. 8.204) as well as the

8.2 Mass Damping Systems Case Studies

wind speed (Fig. 8.205). The data is streamed to a PC via the Internet which controls the sensors, stores several months of data, and each night evaluates the previous day’s acquired data for performance and optimal tuning. This system was used to initially determine that the water depth should be approximately 1 m. Additionally, this system has been used to evaluate the buildings’ performance during typhoons and has verified that the building continues to meet or exceed expectations.

8.2.7.5 Comparison of different design strategies 8.2.7.5.1 Structural options considered Other structural options such as core with outrigger braces were considered at early stages of the design but this option was replaced with distributed concrete dividing walls in response to architectural program.

FIGURE 8.203 Hyatt Park Tower: building accelerometers.

FIGURE 8.204 Hyatt Park Tower: wave sensor.

807

808

CHAPTER 8 Case studies of tall buildings

FIGURE 8.205 Hyatt Park Tower: wind sensor on roof.

8.2.7.5.2 Damping solution considered Both TMD and TLD were considered at early stage. TMD was ruled out due to cost implications.

8.2.7.5.3 Cost
benefit analysis TMD usually occupies smaller areas as opposed to TLD. A cost
benefit analysis was performed comparing the net sellable area gained by using TMD. Considering the dual application of water tank for both fire protection system and added damping, TLD was selected as final option. Damping cost. Cost of dampers as a ratio of total construction cost was approximately 1%.

8.2.7.6 Lesson learned and recommendations 8.2.7.6.1 Difficulties in the design No further information is available.

8.2.7.6.2 Design innovative solutions No further information is available.

8.2.7.6.3 Possible improvements No further information is available.

8.2.8 BLOOMBERG TOWER, NEW YORK CITY, NEW YORK, UNITED STATES 8.2.8.1 Project data The major building data (Fig. 8.206) are summarized as follows:

8.2 Mass Damping Systems Case Studies

FIGURE 8.206 Bloomberg tower: building overview.

• • •

Year of completion: 2004 Developer/contractor: 731 Lexington LP; Vornado Realty Trusy; Bovis Lend Lease (United States) Architectural designer: SLCE Architects (United States); Cesar Pelli & Associates (United States)

809

810

CHAPTER 8 Case studies of tall buildings

• • • • • • • • • • •

Structural designer: Thornton Tomasetti Engineers (United States) Damping supplier: RWDI (design, Canada); Metropolitan Metals (fabrication, United States) Testing laboratory: Wind tunnel tests performed at RWDI Inc. (Canada) Height: 246 m Interstory height: no information available Stories: 55 above grade Gross area of the tower: 130,064 m2 Building function: Condominium, office, retail Structural material: No information available Damping type: TMD Floor plan area: No information available

8.2.8.2 Introduction/History No information is available.

8.2.8.3 Structural system To accommodate the tower’s mixed use, the commercial levels in the lower 30 floors are framed entirely in steel that then support a cast-in-place flat plate structure used for the 24 residential levels above. One of the most critical and challenging aspects of the design was how to transfer the gravity and the lateral loads from the concrete structure above the 30th floor, to the steel frame below. This task was further complicated by two geometry issues: the concrete shear wall was completely misaligned with the steel braced core below, and the building massing had a major setback at the NS elevations so that the steel perimeter columns were outboard of the concrete ones. Both issues prohibited a more direct transfer of the loads. A unique system of outrigger trusses, belt trusses, and transfer trusses was provided between the 29th and 30th floors. The outrigger system consists of four NS and two EW trusses engaging the steel core and the perimeter steel columns. It significantly reduces overturning moment in the core while increasing the building’s lateral stiffness. These outrigger trusses also support a series of transfer trusses and plate girders used to pick up the residential concrete columns. In all, 46 concrete columns were transferred onto the steel frame below. The concrete shear walls were also picked up on secondary steel trusses that were then delivered onto the steel core bracing bays. In order to develop the wall reinforcing into the steel truss elements, as well as achieve good continuity between the two systems, the concrete core was overlapped with the steel trusses by encasing the trusses with concrete and extending the reinforcing within the height of the 30th floor. Reinforcing bars terminated on steel members via welded Lenton couplers. An economical and efficient building emerged from the coordinated use of concrete and steel where each was most suitable. The use of concrete in the residential floors accommodated story heights varying from 3.28 to 4.50 m, resulting in more apartment floors and units within the

8.2 Mass Damping Systems Case Studies

zoning requirement. The typical residential floor structure is a 23-cm thick RC flat plate allowing for easier forming, pouring, and finishing.

8.2.8.3.1 Building fundamental periods The main building fundamental periods were: 6.5 seconds (in X direction) and 6.0 seconds (in Y direction).

8.2.8.3.2 Damping strategy utilized A TMD to satisfy tenant comfort criteria, housed at roof level, has a unique lowheadroom design. This TMD has two linked systems: a steel mass box full of plates on a short pendulum and a second mass box on pivoting legs as an “inverted pendulum.” The linked masses sway out of phase with the tower, creating large differential displacements. Dashpot “shock absorbers” are pushed and pulled to convert kinetic energy into heat. The innovative dual-mass system fits within the 7.6-m vertical height constraint.

8.2.8.3.3 Additional damping provided by the damping system The final TMD weight was approximately 545 tons and the provided total damping was approximately 4.5% of critical. The TMD reduced the building accelerations (for wind storms with a 10-year return period) from 21.1 mg to just below 15 mg.

8.2.8.3.4 Building cost versus damping cost The unique design fits in the shallowest height possible, resulting in additional savings in enclosure costs.

8.2.8.3.5 Building code No information is available.

8.2.8.3.6 Peer-reviewed project No information is available.

8.2.8.3.7 Design forces No information is available.

8.2.8.4 Damping overview 8.2.8.4.1 Damping strategy Wind engineering studies by Rowan Williams Davies & Irwin Inc. (RWDI), demonstrated to structural engineer Thornton Tomasetti, that lateral accelerations on the building’s top levels would be higher than desirable for luxury condominiums. Structural iterations were performed to optimize the structure and reduce the predicted motion. The developer also considered a shorter building: a very effective method of reducing vibration, but one that could critically diminish the viability

811

812

CHAPTER 8 Case studies of tall buildings

FIGURE 8.207 Bloomberg tower: wind tunnel test.

of the development project, since the most valuable real estate is the top portion of the tower. The structural optimization studies performed indicated that due to the multiple structure types and materials, the motion at the top of the building would be difficult to control from a purely structural approach. Since the development program required a less-than-ideal structural system, designers chose to implement a damping system solution to reduce the vibration. Two different types of damping systems were considered: a TMD and a TLCD. Detailed wind-induced structural response studies conducted for the Bloomberg Tower in New York City (Fig. 8.207) predicted 10-year peak acceleration of 22.5 mg. Traditional structural modification measures of adding stiffness and mass were first investigated but the results did not come close to the target performance of 15 mg and, moreover, the proposed structural changes were cost-prohibitive. The remaining remedial measure that could be taken was to install a supplementary damping system that could fit in the available limited space of 7.6-m high. If a simple pendulum-type TMD, like that shown in Fig. 8.208, were to be employed, a floor to ceiling height of greater than 16 m would be required.

8.2.8.4.2 Damping type TMD was chosen, since the space required for TLCD installation in the Bloomberg building was beyond what was available. The TMD weighs 545 tons and in 2004, a uniquely designed articulated mass TMD was installed at the top of the tower (Fig. 8.209).

8.2 Mass Damping Systems Case Studies

FIGURE 8.208 Bloomberg tower: TMD general scheme.

FIGURE 8.209 Bloomberg tower: pictures of installed TMD.

8.2.8.4.3 Structural and damping design This innovative two-mass system can be tuned to long building periods with limited vertical space. The upper 200-ton steel block stands on jointed, destabilizing columns. It is linked to the lower 345-ton steel block suspended from short pendulum cables. As a result, the mass is the sum of the two blocks, but the

813

814

CHAPTER 8 Case studies of tall buildings

equivalent spring stiffness is established by the difference between the negative stiffness effect of the upper mass and the restoring force of the lower mass. The total mass coupled with the low net stiffness results in relatively long periods that can be properly tuned to those of the building. This articulated mass TMD is designed to reduce 10-year peak acceleration to approximately 15 mg by generating a total equivalent damping ratio of 4.5% of critical. Code and guidelines. No information is available. Analysis modeling and software. The wind tunnel data files collected during scale model testing were used as the excitation force in nonlinear numerical simulations, performed by custom RWDI in-house software. Performance and loads were determined resulting from both service level winds and for ultimate design loads (both wind and seismic). Additional numerical simulations were performed with visual Nastran Desktop Motion to verify overall performance for select governing simulations, and to include flexibility of members for more accurate determination of internal and interface loads. Strength-of-materials (stress and strain) calculations were conducted by both fundamental approach (hand calculations) and finite element modeling with COSMOS. Design principles. A TMD at the top of a building moves out of phase as the building sways, driving dashpots, large “shock absorbers” that convert a portion of the kinetic energy of building motion into heat. A passive TMD was selected because it needs no outside energy source, making it reliable and easy to maintain. The simplest TMDs are pendulums whose free length is varied to match the building period, just as the length of the pendulum in an old-fashioned clock is adjusted to better keep time. The natural period of a single degree-of-freedom mass-and-spring system in radians/s is (mass/spring stiffness)0.5. A pendulum is a stable oscillator independent of mass. A pendulum with length L has an equivalent period of (L/g)0.5. So in order to match the tower’s longest periods (6.5 and 7.9 seconds) the pendulum would have to be about 15.5 m long within a space about 18.3 m high. This would require additional costs for framing and cladding the TMD enclosure, and for resisting the greater wind loads created by making the building taller for the additional room. So the design team was challenged to keep the pendulum height to a minimum. The Bloomberg tower TMD uses the principle of linked stabilizing/destabilizing forces to match building periods in a low-headroom space. An inverted pendulum is a “lollipop” balanced on its lower tip. As it leans, it generates a destabilizing force of magnitude exactly opposite to that of a regular pendulum of the same mass and length. Without a restoring force the “lollipop” would flop over and stay there. For large masses, single-point systems are not practical, so a “loose-jointed tabletop” is used instead. If two equal masses, one a pendulum of length L and the other a tabletop with leg length L are linked together, the destabilizing and restoring forces would be equal and opposite. The masses could be pushed sideways and released, and they would stay put. By making the pendulum length a bit shorter than the tabletop, or by making the pendulum heavier than the

8.2 Mass Damping Systems Case Studies

tabletop, a small net restoring force is left to act on the two large linked masses. The resulting system frequency is very low and the period is very long. The resulting innovative two-mass system can be tuned to long building periods but only requires a room 7.6-m high. An upper 200-ton steel block stands on jointed legs. It is linked to a lower 345-ton steel block suspended from short pendulum cables. As a result, the mass is the sum of the two blocks, but the spring stiffness is established by the difference between the destabilizing effect of the upper mass and the restoring force of the lower mass. The high mass and low spring stiffness results in periods long enough to match building behavior. Design phase considerations. Meeting occupant comfort criteria was a constant goal throughout the design process. Very elaborate and detailed parametric studies were carried out in order to determine the effectiveness of increasing the building stiffness by adding structural material, increasing the building’s mass by using thicker floors, and changing the building’s mode shapes by altering its mass distribution with height through varying slab thicknesses. These options were available because wind governed the lateral design of the main tower; the seismic requirements were moderate. However, wind tunnel model testing at Rowan Williams Davies & Irwin, Inc. revealed that local conditions create a power spectrum, or graph of wind energy versus the building’s period, that is flat in the range of lowest-mode periods. This meant that modifying the tower’s dynamic properties by changing stiffness, mass, or mode shape would do little to affect the building’s response and occupant comfort levels in windy conditions. Some changes could even worsen the response. In addition, any of these modifications would be cost-prohibitive. Two possible options for providing additional damping to the structure were studied: a TMD or a TLCD. After careful consideration of advantages and disadvantages for both systems, such as cost, size, future adjustability, and more, a passive TMD system was selected by the owner and the design team. Once the system was selected, the design effort was then focused on how to optimize the TMD and minimize the height of the pendulum.

8.2.8.4.4 Architectural integration strategy The TMD is entirely contained within a single room atop the building. Very little architectural revision was required to accommodate this room.

8.2.8.4.5 MEP integration strategy No information is available.

8.2.8.4.6 Elevator and other mechanical devices integration strategy No information is available.

8.2.8.4.7 Experimental tests No information is available.

815

816

CHAPTER 8 Case studies of tall buildings

8.2.8.4.8 Monitoring system Full-scale measurements of the dynamic response of the building are taken. During the measurements period, the building is excited by ambient winds approaching the site from a variety of directions. In addition to winds, the operation of the tower crane also results in building oscillations. A total of four accelerometers were utilized in the measurements. Two accelerometers were used to measure NS motion (Y direction) and another two accelerometers were used to measure EW motion (X direction). The twist (torsion) was estimated from taking the difference between the signal of any two accelerometers pointing in the same direction. Several 20-minute-long records of the accelerometers signals were taken. A typical time history that was recorded is depicted in Fig. 8.210. In order to determine the frequencies of the time history data that were collected, a mathematical algorithm (fast Fourier transform (FFT)) was used. For each individual accelerometer, the result of this algorithm is an autospectrum, and an example of these is presented in Fig. 8.211. Peaks in the FFT magnitude plots, in association with the broadband excitation mechanism of ambient wind excitation, indicate the natural frequencies of the structure. The comparison between the FFT phase plots allows us to distinguish between sway and twist modes. Each set of accelerometers pointing in the same direction will have zero phase difference at the sway mode and 180-degree phase difference at the twist mode. A peak at frequency of 0.265 Hz is attributed to crane movement and not to any fundamental frequency of the building.

FIGURE 8.210 Bloomberg tower: typical sample of time history acceleration measurements.

8.2 Mass Damping Systems Case Studies

300

FFT-x1 (magnitude)

250 200 150 100 50 0 0

0.1

0.2

0.3

0.2

0.3

0.4 Frequency (Hz)

0.5

0.6

0.7

0.8

0.5

0.6

0.7

0.8

200 150

FFT-x1 (phase)

100 50 0 –50

–100 –150 –200 0

0.1

0.4 Frequency (Hz)

FIGURE 8.211 Typical FFT magnitude and phase of x1 acceleration measurement.

8.2.8.4.9 Maintenance strategy Cursory glances into the room by building maintenance staff are integrated in the preventive maintenance plan. Annual visits have been made by RWDI to inspect for indications of functional performance (travel indicators, alignment/index marks, etc.) and cosmetics (paint flaking, rust, etc.). After the first couple of visits, a very short punch list of outstanding items has been cleaned up, and nothing of interest has been noted in subsequent years. Regular RWDI inspections also check for continued weld and bolt integrity. After 5 years, the cable tensions were checked to confirm that they retain equal load sharing.

8.2.8.5 Comparison of different design strategies 8.2.8.5.1 Structural options considered No information is available.

817

818

CHAPTER 8 Case studies of tall buildings

8.2.8.5.2 Damping solution considered TLCDs were familiar to the structural design team, and favored for their low cost. However, sufficient damping capacity could not be integrated into the available space. The opposed-pendulum TMD design was therefore developed to fit into the short available vertical space instead.

8.2.8.5.3 Cost
benefit analysis No information is available.

8.2.8.6 Lesson learned and recommendations 8.2.8.6.1 Difficulties in the design No information is available.

8.2.8.6.2 Design innovative solutions This was the first large opposed damper deployed. Anticipating that the as-built frequencies of the tower would be dissimilar to those predicted by computer model, a frame for adjusting the TMD pendulum frequency was designed so that it could be properly tuned to the as-built structural frequencies.

8.2.8.6.3 Possible improvements No information is available.

8.2.9 RAFFLES CITY, CHONGQING, CHINA 8.2.9.1 Project data The major building data (Fig. 8.212) are summarized as follows: • • • • • •

• • • •

•

Year of completion: 2018 (under construction) Developer/contractor: Capitaland Investment Co. Ltd (China) Architectural designer: Moshe Safdie & Associates Ltd (China) Structural designer: Arup (Shanghai and International) Damping supplier: Taylor Devices Inc. (United States) Testing laboratory Wind: RWDI (Canada) Seismic: Tongji University (China) Height: T2/T5/T3S/T4S 239 m (from bottom slab) Interstory height: T2/T5 3.5 m; T3S 3.4 m; T4S 4.3 m Stories: T2/T5 48 levels; T3S 50 levels; T4S 54 levels (above podium); 6 levels podium and 3 levels basement Gross area of the tower: T2/T5: 73,000 m2 each T3S/T4S: 77,000 m2 each Conservatory: approximately 9000 m2

8.2 Mass Damping Systems Case Studies

FIGURE 8.212 Raffles City: building overview.

• •

•

•

Building function: T2/T5: residential; T3S: commercial/office; T4S: office; Conservatory: leisure/restaurant Structural material Concrete classes (Mainland China); piles: C40; foundation mat: C40; columns: C50/C60; cores: C50/C60; slab: C35/C40; belt truss/outrigger truss: C40 (structural steel embedded) Structural steel (Mainland China): composite columns: Q345B; belt truss: Q345B; outrigger truss: Q345B; conservatory structural truss: Q345B; rebar: HRB400/HRB500 Damping system: Six friction pendulum bearing (FPB) in each tower expect T4s that has 8 of them Floor plan area: No information is available.

8.2.9.2 Introduction/history No information is available.

8.2.9.3 Structural system T2/T5/T3S/T4S are typically RC construction with steel RC (SRC) columns. The signature curvature of the towers is carried through the structure with architecturally expressed curving columns at the fac¸ade. Towers carry the north curvature of

819

820

CHAPTER 8 Case studies of tall buildings

the tower through the podium level to create a curved atrium space within the podium. All other columns become vertical at the top of podium level. Curving slots are introduced into the fac¸ade to allow light and air into the residential units. The slots follow the curvature of the columns. The central core provides the primary lateral support and is positioned to allow openings for the curving slot. The lateral force resisting system of the tower consists of an RC central core that engages an RC and SRC perimeter frame with outriggers and belt trusses. The core serves as the primary lateral load resisting system, while the outriggers and perimeter frame serve as the secondary system. The slenderness ratio of T2/T5 and T3S/T4S are 8.35 and 7.47, respectively, while the core height:width ratio of T2/T5 and T3S/T4S are 18.13 and 23.53, respectively. The central core engages the curved perimeter with one level of outriggers placed at MEP/refugee floors for T2/ T5, and three levels of outriggers placed at MEP/refugee floors for T3S/T4S. Conservatory structural system: The primary structure comprises three longitudinal steel trusses spanning EW between the supporting towers. The trusses are designed as continuous spans over the buildings. Regularly spaced transverse trusses connect the three primary trusses together; 250-mm thick slab-on-metaldeck spans approximately 4.5 m between the transverse trusses while a perimeter edge beam receives the roof structure above.

8.2.9.3.1 Building fundamental periods The building fundamental vibration modes are shown in Fig. 8.213. Two different structural software have been utilized and the comparison between the different modal properties are the following: • •

SAP 2000: 1st mode: 6.09 seconds; 2nd mode: 4.12 seconds; 3rd mode: 4.01 seconds LS-DYNA 2000: 1st mode: 6.18 seconds; 2nd mode: 4.25 seconds; 3rd mode: 4.23 seconds

FIGURE 8.213 Raffles City: fundamental frequencies and mode shapes.

8.2 Mass Damping Systems Case Studies

8.2.9.3.2 Damping strategy utilized Chongqing Raffles city uses different types of bearings to connect the tower and conservatory which include fixed (full fixed or partial fixed), LRB (lead rubber bearing), FPB and FPB combined with damper. There are six FPBs in the each tower except eight in T4s tower and two perpendicular dampers in each tower (Fig. 8.214). This scheme can reduce the bridge displacement and internal force in the tower and the comfort can also be improved.

8.2.9.3.3 Additional damping provided by the damping system The modulus of the structure is mainly influenced by the four tower. The damper will not change the modal characteristics too much. Comparing with the fixed method, the seismic bearing scheme only reduced about 3%.

8.2.9.3.4 Building cost versus damping cost Reduction of the seismic forces in the bridge truss and consequent saving in steel tonnage and construction cost.

8.2.9.3.5 Building code The major standard Chinese design codes used in the projects are the following: • • • • •

Standard for classification of seismic protection of buildings (GB 50223, 2008) Load code for the design of building structures (GB 50009, 2012) Code for seismic design of buildings (GB 50011, 2010) Code for design of concrete structures (GB50010, 2010) Code for design of steel structures (GB50017, 2003)

FIGURE 8.214 Raffles City: bearings location.

821

822

CHAPTER 8 Case studies of tall buildings

• • • • •

Technical specification for steel structure of tall buildings (JGJ 99, 1998) Technical specification for tall RC buildings (JGJ 3, 2010) Technical specification for box and raft foundation for tall buildings (JGJ 6, 2011) Technical specification for steel RC composite structures (JGJ 138, 2001) Specification for design and construction of concrete-filled steel tubular structures (GB50936, 2014)

8.2.9.3.6 Peer-reviewed project In the case of this extremely irregular buildings with Sky Bridge, the design team is required to have a National Expert Panel Review, consisting of a number of eminent academics, engineers, and researchers. The project team passed this rigorous review and gained special approval from the experts on August 8, 2014.

8.2.9.4 Damping overview 8.2.9.4.1 Damping strategy The final articulation scheme is illustrated in Fig. 8.215. This positioning takes into account the latest geometry of the conservatory support truss and the location of electrical and mechanical plants at the tower interface. It also aims at facilitating maintenance access and reducing congestion while maintaining isolation efficiency. Considering the space in the top between the tower and conservatory is limited, which cannot install too many bearings on the towers. So the dampers can not only reduce the displacement but also save the space for structure.

8.2.9.4.2 Damping type Viscous dampers have been utilized at the top of the building. The relative properties are given in Table 8.40.

8.2.9.4.3 Structural and damping design Code and guidelines. Code for seismic design of buildings • •

Seismic isolation Damper application

FIGURE 8.215 Raffles City: damping articulation scheme.

8.2 Mass Damping Systems Case Studies

Analysis modeling and software. Due to the highly nonlinear behavior of the FPBs and the possible phase differences between the tower displacements, response-spectrum analysis methods are not appropriate. Instead, the seismic response of the four supporting towers and the conservatory bridge structure to the level 3 MCE was simulated in LS-DYNA by nonlinear time history analysis method. Seven sets (five natural and two artificial records) of spectrum-compatible ground motion time histories were applied for nonlinear transient analysis in the detailed design and validation stages. Towers and conservatory models Early assessment models included elastic representations of the towers and the conservatory. But for the isolation scheme detailed design and verification stages, fully nonlinear models were developed to properly capture member damage, hysteretic energy dissipation, and structural period elongation effects. The design of the structure has been constantly updated as the project advanced. Details of the conservatory/tower interface were introduced in the analysis models as design choices were being finalized. For instance, the location and size of the elevator shafts, fire escapes, and the detailed structure of the pedestrian link bridge to the north towers imposed additional constraints on the articulation scheme. Assessment of seismic demands and damage of structural members. The same comprehensive model has been used to assess the seismic performance of the conservatory and of the supporting towers. The complete analysis model is illustrated in Fig. 8.216. Interface elements, FPB, viscous dampers are modeled explicitly (Fig. 8.217). Connection detailed design and locations have been updated as the structural envelope was progressively frozen. Design principles. No information is available. Design phase considerations. At the first stage, the bridge is fixed on the tower which causes larger shear force in the rare earthquake case. So the design team decide to use flexibility connection. LRB, FPB, and bearing combined with damper were considered with different sizes and parameters. For the pure LRB scheme, vast bearings will be needed. But comparing it with the fixed model the shear force reduces about 61%. After adding the dampers into the model, the displacement reduces about 150% compared with pure FPB scheme. So the FPB can reduce the friction coefficient which is much more stable in high-speed situation. The shear can also be reduced although it may not be as effective as displacement, but it is beneficial to the design.

8.2.9.4.4 Architectural integration strategy No information is available.

8.2.9.4.5 MEP integration strategy No information is available.

823

824

CHAPTER 8 Case studies of tall buildings

FIGURE 8.216 Raffles City: complete LS-DYNA analysis mode.

8.2.9.4.6 Elevator and other mechanical devices integration strategy No information is available.

8.2.9.4.7 Experimental tests No information is available.

8.2.9.4.8 Monitoring system No information is available.

8.2.9.5 Comparison of different design strategies 8.2.9.5.1 Structural options considered Seismic isolation of the conservatory bridge can be achieved by means of elastomeric or sliding bearings. Realistic modeling parameters and provisional cost estimates were obtained from established suppliers: DIS, SRIM, for LRB solutions, from Mageba, Maurer Soehne GmbH, SRIM, for FPBs, from Maurer Soehne, Taylor Devices, LuoYang ShuangRui for viscous dampers.

8.2 Mass Damping Systems Case Studies

FIGURE 8.217 Raffles City: details of conservatory/tower interface
final scheme.

LRBs were considered at concept stage but later on abandoned: the reduced bearing capacity at large shear displacements and the lower shear resistive force compared to large FPB proved that LRBs are prohibitive. Options combining FPBs and viscous dampers were also assessed. The main contribution of the viscous dampers, apart from increasing the damping in the structure, is to reduce the displacements at the conservatory/tower interface in both longitudinal and transverse directions.

8.2.9.5.2 Damping solution considered The final scheme is FPB combined with dampers. The scheme was optimized with different velocity exponents of damper.

8.2.9.5.3 Cost
benefit analysis When the conservatory is isolated at all towers, the reduction of peak resultant shear forces at the interface may exceed 200% enabling significant material savings and design simplification of the supporting structure at the top of the towers. The pure FPB scheme can save about 400-ton steel comparing with the case the

825

826

CHAPTER 8 Case studies of tall buildings

conservatory is fixed, and in the case FPB 1 damper is used there will be even more savings. Damping cost. The predicted steel tonnage saving on the conservatory alone is estimated to 1322 ton. Table 8.38 provides the details of the assessment. Design Implication. Table 8.39 summaries the different key points with and without the use of damping: structural element size/reinforcement ratio/acceleration/wind and seismic forces.

8.2.9.6 Lesson learned and recommendations 8.2.9.6.1 Difficulties in the design The following difficulties were encountered during the design: (1) the limitation of the code and specification, the need to collaborate with seismic experts to determine its design performance criteria; (2) multitower analysis under different loading cases (wind, seismic); and (3) reasonably determine the size of FPB and damper’s performance.

8.2.9.6.2 Design innovative solutions The following innovative solutions were developed during the design: Setting different performance criteria for dampers and FPBs under different serviceability and limit states. Increase the redundancy of energy dissipation in high levels by mixing using FPBs and dampers, coupling with reducing the size of FPBs, saving construction complexity and cost.

8.2.9.6.3 Possible improvements No information is available. Table 8.38 Steel Tonnage Reduction—Conservatory Structure Structural Member

Bidding Phase (t)

Optimization (t)

Saving (t)

Optimize Method

Primary truss Secondary trusses Plane supporting Small bridge Bearing components Lift slab and floor beam Total

2822 1425

2200 1325

622 100

Isolation bearing Stiffness optimize

1042

692

350

453 623

453 623

0 0

Stiffness/slab optimization at MCE



500

250

250

6865

5543

1322

Modification on primary truss


Table 8.39 Advantages of Damper Using Damper 1 FPB

FPB Only

• Little impact on the displacement because of weight • Movement joint around elevator shafts and bridge seem manageable • Integrated fuse function (low maintenance costs after the earthquake) • Lower speed index can guarantee lock fixed on the normal wind conditions

• Great impact on the displacement because of weight and friction • Currently no appropriate technical solution for the movement joint • Require specific mechanical fuse function (need to be replaced after the earthquake)

• Force distribution is more reasonable • Fatigue load is relatively small

• More accurate releasing force • Can adapt thermal displacement and deformation during the construction phase • Do not produce out of situation on the actual situation • Do not need to set the tension mount • Using more commonly used dynamic friction coefficient 5% • More commonly used performance parameters

• Easily reset

• To resist wind load of normal conditions by friction is not recommended in the specification (the unpredictable vibration can make friction coefficient decreases as the value of dynamic friction coefficient) • Failure time of different tower fuse is not the same • Fatigue load under the joint action of temperature and wind loads is a greater impact on the fuse • To estimate the releasing force is difficult • Because gravity and temperature loads will cause prestress, set fuse after even bridge construction • The most serious conditions have 45 mm detachment-need to set the tension mount • Higher dynamic friction coefficient 7% • Difficult to maintain a stable state • Having a risk of viscosity—slip effect, and then leading a high acceleration of even bridge • Having a potential problems of returning to the center

Table 8.40 Raffles City: Viscous Dampers Design Parameters Design Parameters

Value 5 Optimized Scheme

Constitutive law

F 5 CVα F: Force (N) V: Velocity (m/s) α: Velocity exponent C: Constant (Ns/m)2α α 5 0.3 C 5 5E6 MN/(m/s)0.3 tuned to obtain 5000 kN at V 5 1 m/s 4 mm per tower (2 in longitudinal and 2 in transverse direction) 400 mm at towers T2, T3s, and T5 450 mm at tower T4S 1E6 cycles at 900 kN service wind load Fuse release force 2 MN at approximately 1.5 mm/s

Law parameters Configuration Maximum stroke Fatigue resistance Integrated fuse function

828

CHAPTER 8 Case studies of tall buildings

8.2.10 L-TOWER, TORONTO, CANADA 8.2.10.1 Project data The major building data (Fig. 8.218) are summarized as follows: • • • • • • • • • • • •

• •

Year of completion: 2015 Developer/contractor: Canderel Residential (Canada) Architectural designer: Studio Daniel Libeskind (United States), Page 1 Steele/IBI Group (Canada) Structural designer: Jablonsky Ast & Partners (Canada) Damping supplier: Gradient Wind Engineering Inc. (Canada) Testing laboratory: Gradient Wind Engineering Inc. (Canada) Height: 234 m Interstory height: 2.95 m Stories: 58 above grade Gross area of the tower: 38,814 m2 (417,700 ft2) Building function: Condominium Residential/Retail/Performing Arts Structural material Concrete classes: Concrete strength at 28 days 25 MPa (top) to 55 MPa (base) Structural steel: NA Dampers: not used Rebar: 400 MPa Floor plan area: trapezoid with dimension of 31 m 3 33 m for a floor of 1010 m2 Structural system: ductile concrete core shear wall with column capitals

8.2.10.2 Introduction/history The tower is 58-story tower constructed at the southwest corner of a redevelopment site. The tower and the podium of the development forms an L-shape with a curved fac¸ade at the north elevation. The fac¸ade on north elevation has slight over hang above the podium level up to approximately two-thirds of the tower height.

8.2.10.3 Structural system The structure adopts a ductile concrete core shear wall with column capitals (Figs. 8.219 and 8.220).

8.2.10.3.1 Building fundamental periods The natural vibration periods of the tower were estimated to be 4.9 seconds along the longer direction of building plan and 4.1 seconds in shorter direction and 3.8 seconds in torsion. The three lowest vibration modes are illustrated in Fig. 8.221. Definition of coordinate axes is illustrated in Fig. 8.222. Average mass density of top floors is approximately 312 kg/m3, which is equivalent to weight density of 19.5 lb/ft3.

8.2 Mass Damping Systems Case Studies

FIGURE 8.218 L-Tower: building overview. © Edward Chamberlain via CTBUH.

829

830

CHAPTER 8 Case studies of tall buildings

FIGURE 8.219 L-Tower: typical floor plan.

FIGURE 8.220 L-Tower: elevations.

8.2 Mass Damping Systems Case Studies

Mode 1 (4.9 s)

Mode 2 (4.1 s) 250

250

Mode 3 (3.8 s) 250

200

200

150

150

150

100

100

100

50

50

50

Height (m)

200

–0.50

–0.40

–0.30 –0.20 –0.10 Ux Uy R2*Rg

Mode 1 : T1 = 4.9 s

0

0.00

–1.00

0

–0.50 Ux

0.00 Uy

0 0.50 R2*Rg

Mode 2 : T2 = 4.1 s

FIGURE 8.221 L-Tower: fundamental frequencies and mode shapes.

FIGURE 8.222 L-Tower: definition of coordination axis.

1.00

–1.00

–0.50 Ux

0.00 Uy

0.50 R2*Rg

1.00

Mode 3 : T3 = 3.8 s

831

832

CHAPTER 8 Case studies of tall buildings

8.2.10.3.2 Damping strategy utilized Wind tunnel tests predicted that the building will experience high-amplitude vibration in both lateral directions. Additional damping of 1.7% and 2.3% were required to reduce the accelerations to the acceptable level based on industry guidelines. Since the accelerations in two translational directions were similarly high, a two-directional TLSDs were designed on level 60.

8.2.10.3.3 Additional damping provided by the damping system Additional damping values introduced by the TLSD were approximately 4.5% and 3.5% in first and second modal directions, respectively, which satisfies the additional damping requirement. The additional damping value of the damper has been verified through a 1:10 scale dynamic building-damper testing in dynamic test rig facility at Gradient Wind Engineering Inc.

8.2.10.3.4 Building cost versus damping cost Estimated damping cost: 1 million in CAD (almost US$800,000).

8.2.10.3.5 Building code The following is a list of the major standard design codes used in the project: • • • • • • •

CSA S16-09—design of steel structures (CSA, 2009) CSA 23.3-04—design of concrete structures (CSA, 2004) Ontario Building Code, 2006, 2012 (OBC, 2006, 2012) National Building Code of Canada 2005, 2010 (NBCC, 2005, 2010) ASCE 7-05 (ASCE, 2005) Eurocode 1 (CEN, 2010) International Standard Organization (ISO): ISO-6897 (ISO, 1984), ISO-10137 (ISO, 2007)

8.2.10.3.6 Peer-reviewed project None.

8.2.10.3.7 Design forces Wind loads were determined based on statistical model of the wind speed and direction data measured at Pearson International Airport in Toronto for the last 40 years. The measured wind speeds were calibrated to give a 50-year wind speed of 28.4 m/s at 10 m above grade corresponding to a dynamic pressure of 0.52 kPa, and 10-year dynamic pressure of 0.39 kPa.

8.2.10.3.8 Expected performance The 1- and 10-year return period peak wind-induced accelerations at 58th floor, the top occupied floor, were predicted to be 21.1 and 21.2 mg, respectively. Peak torsional velocity was predicted to be small. Based on guidelines of ISO, the

8.2 Mass Damping Systems Case Studies

building acceleration exceeded the recommended maximum 10-year return period acceleration of 18 mg, which was not acceptable.

8.2.10.4 Damping overview To mitigate the acceleration levels below the serviceability limit of 18 mg, the building requires approximately 2.3% of additional damping for both Y and X directions in addition to the inherent structural damping.

8.2.10.4.1 Damping strategy Building accelerations were measured to be higher than the industry guidelines (ISO 6897 (ISO, 1984) and ISO 10137 (ISO, 2007)) based on a wind tunnel testing, which was also confirmed by in situ building monitoring. Among supplemental damping systems including TMD and TLSD, TLSD system is used for the building due mainly to the cost-efficiency. Preliminary TLSD was designed based on dynamic structural properties of computer model. Preliminary tank dimensions were used for tank construction considering some rooms which were to be determined pending outcome of in situ building monitoring. By adjusting the location of solid partitions, it was possible to properly tune the water sloshing frequency to the as-built building frequency which is increased and reduced from the computer model. The main interest of the design was on minimizing the space especially in plan, while achieving required supplemental damping. To minimize the main potential problem of water leakage, screen posts were designed to be anchored on the ceiling (Fig. 8.223). In final design of the damper, since the as-built building periods were measured to be shorter than those of theoretical model by approximately 15%
20%, for the tuning purpose, the water depth was increased to 3.5 m from the original depth of 3.22 m, and the effective length of the tank was reduced by relocating the solid screens (Fig. 8.228). The damper designed to satisfy the guidelines for building serviceability design, however, was not utilized for the strength design of the tower in compliance with current design practice.

8.2.10.4.2 Damping type As shown in Fig. 8.228, a bidirectional TLSD is designed on level 60 to mitigate the building vibration to the acceptable level. To get the required damping, the damper was designed to have effective water mass ratio (equal to the ratio of effective water mass to the building modal mass) of 0.9% and 1.0% in X and Y-directions, respectively. The sloshing frequency in each direction was tuned to its optimal target frequency which is similar to the natural frequency of the building. To generate optimal damping to dissipate the maximum vibration energy, two 45% porous internal baffles were installed at approximately 0.4L and 0.6L locations along the length of the tank in both directions, where L indicates tank length.

833

834

CHAPTER 8 Case studies of tall buildings

FIGURE 8.223 L-Tower: schematic drawing of bidirectional TLSD installed at level 60.

8.2.10.4.3 Structural and damping design Code and guidelines. The damper was designed to provide required damping under the wind loads measured from 1:400 scale HFFB measurement of the building in wind tunnel as shown in Fig. 8.224. The 10-year return period hourly wind pressure of 0.39 kPa were used for the test based on NBC (2005)/OBC (2006); wind tunnel tests are conducted to meet or exceed the requirements of “Wind Tunnel Studies of Buildings and Structures,” ASCE Manual 7 Engineering Practice Note 67 (ASCE, 1999). Analysis modeling and software. Before the dynamic test, a linear time history analysis of sloshing damper-building coupled system was performed to predict the performance of the damper in both X and Y directions. Design parameters such as spring stiffness and mass as well as the time history of actuator forces were verified and determined from the analysis. In the analysis, a scaled building model and a damper was modeled as a two degree-of-freedom system in model scale. Wind loads measured from a wind tunnel test (Fig. 8.229) were scaled in magnitude and time, and applied to the model for the time-domain analysis. In

8.2 Mass Damping Systems Case Studies

FIGURE 8.224 L-Tower: wind tunnel test setup.

the analysis, an in-house Matlab code was used. In the code, TLSD was modeled as a single degree-of-freedom equivalent secondary mass whose equivalent mass and screen damping forces were estimated based on analytical solution of potential flow (Tait, 2008). Design principles. The damper was designed to reduce the 10-year peak sway building acceleration at the top of the building of 17 and 22 mg to 13 and 15 mg in X and Y directions, respectively. The additional damping required to achieve the target acceleration is estimated to be approximately 1.7%
2.3% of critical which is to be added to the inherent structural damping of 1.5 % of critical. The damper is tuned to the two lowest lateral vibration modes. Based on the required supplemental damping in X and Y directions, design parameters such as (effective) water mass ratios, water sloshing frequencies, and screen damping were evaluated to achieve the optimal damping. A water height was assumed and length of the tank was determined to get the target sloshing frequency based on the potential flow theory. Tank width was determined to achieve the required mass ratio. Since the length in one direction corresponds to the width in the other direction in bidirectional tanks, the two dimensions were determined iteratively to satisfy both the frequency and mass ratio by adjusting the water height. Design phase considerations. At the preliminary design stage, preliminary dimensions of the tanks and water height were estimated. To verify the additional damping, a dynamic damper/building model test was performed in 1:10 scale under dynamic wind loads measured in wind tunnel. The structural and architectural designs were conducted to ensure the structural safety of supporting structures and walls. From a long-term field monitoring of building motion, the as-built building properties such as natural frequencies were measured and the

835

CHAPTER 8 Case studies of tall buildings

inherent damping of the structure was inferred. From the long-term field measurement before the installation of the damper, it was found that the as-built building frequencies were higher than those from numerical model; accelerations of the building were similar to the wind tunnel measurement; and inherent structural damping was around 1.5%
1.75% to the critical. By using the building frequencies measured on-site, the damper frequencies have been retuned. Documents for the damper design include detailed drawings and waterproofing specifications. See Fig. 8.225 for some detailed drawings.

8.2.10.4.4 Architectural integration strategy The damper tank is located at the 60th level of the building. Due to building’s unique curved profile, the tank placement and depth in elevation provided design 11500 1700

1700

1700

1700

1400

650

650 2000

1800

Screen (porosity 45%)

Interior solid partitions Screen (porosity 45%)

836

1900

2000 Screen (porosity 40%) 12500

A

A

2200

Screen (porosity 40%)

Y

Typical stainless steel column anchored to U/S of concrete cover

2200

2200

X

Plan of TLSD (top removed) FIGURE 8.225 L-Tower: detail damper drawing.

8.2 Mass Damping Systems Case Studies

challenges for the team. To accommodate the tank, the floor height has been adjusted at the tank location. By leaving the space below the tank as a nonresidential and mechanical space, water leakage directly to the residential units are avoided in the design.

8.2.10.4.5 MEP integration strategy Water supply and drain plumbing are integrated to the tank to support regular maintenance. As the tank is interior to the building no heating is required.

8.2.10.4.6 Elevator and other mechanical devices integration strategy Information is not available.

8.2.10.4.7 Experimental tests The performance of sloshing damper was verified through two series of 1:10 scale dynamic damper tests in both X and Y directions. Figs. 8.226 and 8.227 show images of dynamic test rig and 1:10 scale sloshing water tank used for the verification test. A pendulum mass and water tank represent approximately 25% of 1:10 scale building modal mass and TLSD, respectively. The mass designed to move freely in pendulum motion was connected to springs modeling building stiffness. The system was instrumented with accelerometers and load cells to measure the acceleration of the mass, actuator excitation force, as well as the sloshing water force on the screens and tank end walls. To simulate the wind excitation force, wind load time history measured from HFFB test of the building was used. From the test, sloshing water frequency, screen damping, as well as supplemental damping of the damper were measured (Fig. 8.227). For the water

FIGURE 8.226 L-Tower: overall view of pendulum-type dynamic test rig with a scaled TLSD tank.

837

838

CHAPTER 8 Case studies of tall buildings

FIGURE 8.227 L-Tower: close-up view of TLSD tank model equipped with load cell to measure the dynamic water force on internal screens.

depth-to-length ratio, theoretical sloshing frequencies based on potential flow theory matched very well with the test results. To find the optimal porosity of internal screens which can provide the highest damping, tests were performed for a range of screen porosities. Lastly, supplemental damping introduced by the damper were measured to be compared to the required damping.

8.2.10.4.8 Monitoring system Long-term full-scale building monitoring was conducted to measure the as-built building frequencies and damping as well as building acceleration during strong wind storms. The building was instrumented with three accelerometers to measure two translational and torsional accelerations at the damper floor. From the measurement, it was found that the as-built building frequencies are higher than corresponding theoretical values by approximately 15%
20%, which was shown in other buildings. The monitoring will continue after the installation of the damper to verify the efficiency of the damper during the strong wind events.

8.2.10.4.9 Maintenance strategy Due to its passive nature, TSDs have lower maintenance requirements as compared to other systems. Principle maintenance tasks consist of two parts, keeping the proper water level and ensuring waterproofing during the lifetime of the damper. To maintain the water level, regular visual inspection through a transparent vertical glass or automatic water level controller with automatic water filling system has been used. Regular inspection is planned to avoid leaking. The tank

8.2 Mass Damping Systems Case Studies

will be drained every 2 years for a direct inspection of the waterproofing liner. The screen components of the tank are corrosion-resistant stainless steel columns with structural plastic screens. The tank is refilled to the water level within a couple of days after the complete drainage.

8.2.10.4.10 Development process Based on wind tunnel testing and preliminary damper drawing, preliminary sloshing damper designs were prepared by wind laboratory followed by structural design of the supporting structures and slabs. Detailed drawings prepared by the wind laboratory are reviewed by the design office before the installation of the damper.

8.2.10.5 Comparison of different design strategies 8.2.10.5.1 Structural options considered No significantly different structural options were considered but only refinements of the same system described previously, such as the use of transfer slab thicknesses, outrigger system at rooftop.

8.2.10.5.2 Damping solution considered Different types of damping were considered including TMD. However, in terms of cost-efficiency and constructability, TLSD was considered to be the better solution.

8.2.10.5.3 Cost
benefit analysis Damping cost. Estimated damper costs 1.0 million $CDN (almost US$800,000).

8.2.10.6 Lesson learned and recommendations 8.2.10.6.1 Difficulties in the design The large size of the damper tank and its integration with the building vertical profile was a challenge. As described earlier, the damper tank is located at the 60th level of the building. Due to building’s unique curved profile, the tank placement and depth in elevation provided design challenges for the team. As shown in the section drawing in section (Section 8.2.10.4.4) “Architectural Integration Strategy,” to accommodate the tank, the floor height has been adjusted at the tank location. By leaving the space below the tank as a nonresidential and mechanical space, water leakage directly to the residential units is avoided in the design.

8.2.10.6.2 Design innovative solutions No information is available.

8.2.10.6.3 Possible improvements No information is available.

839

840

CHAPTER 8 Case studies of tall buildings

8.2.11 14 YORK STREET, TORONTO, CANADA 8.2.11.1 Project data The major building data (Fig. 8.228) are summarized as follows: • •

• • •

Year of completion: 2015 Developer/contractor: Lanterra developments (Canada) Cadillac Fairview/ Lanterra Construction Management (Canada) Architectural designer: Architects Alliance (Canada) Structural designer: Jablonsky, AST and Partners (Canada) Damping supplier: Gradient Wind Engineering Inc. (Canada)

FIGURE 8.228 14 York Street: building overview.

8.2 Mass Damping Systems Case Studies

• • • • • • •

Testing laboratory: Gradient Wind Engineering Inc. (Canada) Height: 234 m Stories: 67 above grade Gross area of the tower: approximately 58,000 m2 (630,000 ft2) Building function: condominium residential/retail Structural material: No information available Floor plan area: No information available

8.2.11.2 Introduction/history No information is available.

8.2.11.3 Structural system The plan of the tower has a curvilinear square shape with ductile concrete core shear wall with flat slabs (Figs. 8.229 and 8.230).

FIGURE 8.229 14 York Street: typical floor plan.

841

842

CHAPTER 8 Case studies of tall buildings

FIGURE 8.230 14 York Street: elevation.

8.2.11.3.1 Building fundamental periods The natural frequencies of the tower are 0.149 and 0.195 Hz in translational directions and 0.298 Hz in torsional direction which corresponds to 6.7, 5.1, and 3.4 seconds in building periods. Average mass density (per volume) of the building is approximately 390 kg/m3, which corresponds to weight density of 24 lb/ft3.

8.2.11.3.2 Damping strategy utilized Wind tunnel tests predicted that the building will experience high-amplitude vibration in both lateral directions. Additional damping of 3.2% was required to reduce the accelerations to the acceptable level based on industry guidelines.

8.2 Mass Damping Systems Case Studies

Since the accelerations in two translational directions were similarly high, a bidirectional TLSD was designed for the 67th mechanical floor of the building to get best efficiency from the damper.

8.2.11.3.3 Additional damping provided by the damping system Additional damping introduced by the TLSD were approximately 3.5% and 4% in X and Y directions, respectively. The additional damping values of the damper were measured through a 1:10 scale dynamic building-damper test in dynamic test rig facility at Gradient Wind Engineering Inc.

8.2.11.3.4 Building cost versus damping cost An estimate of total damper cost of structure 1 internal equipment including mechanical is 1 million CDN (US$800,000).

8.2.11.3.5 Building code Following is a list of the major standard design codes used in the project: • • • •

CSA S16-09—design of steel structures (CSA, 2009) CSA 23.3-04—design of concrete structures (CSA, 2004) National Building Code of Canada 2005, 2010 (NBCC, 2005, 2010) International Standard Organization (ISO, 2007)

8.2.11.3.6 Peer-reviewed project None.

8.2.11.3.7 Design forces Wind loads were determined based on statistical model of the meteorological data (wind speed and direction data) measured in Pearson International Airport in Toronto over 40 years. The measured wind speeds were calibrated to give a 50year return period wind speed of 29.4 m/s at 10 m above grade corresponding to a dynamic pressure of 0.52 kPa, and 10-year return period wind speed of 25.5 m/s and corresponding dynamic pressure of 0.39 kPa as specified in Ontario Building Code (OBC, 2006).

8.2.11.3.8 Expected performance The 1- and 10-year return period peak wind-induced accelerations on 67th floor, the top occupied floor, were predicted to be in a range between 24 and 30 mg depending on massing configurations tested. Torsional velocity is predicted to be low. The predicted 10-year peak accelerations were well in excess of the target of 18 mg derived from consideration of the International Standards Organization (ISO, 2007) recommendations.

843

844

CHAPTER 8 Case studies of tall buildings

8.2.11.4 Damping overview To mitigate the acceleration levels below the serviceability limit of 18 mg, the building requires approximately 3.2% of additional damping in both Y and X directions in addition to the inherent structural damping.

8.2.11.4.1 Damping strategy The following requirements are considered in damper design: (1) building interstory drift and corresponding deflection should be minimized, therefore the mechanism of the damper should be efficient with small excitation amplitude; (2) since the building will experience similarly high level of accelerations in both translational directions, damping mechanism which can be easily applied to both directions is preferred; (3) the damper should be economical and easy to maintain; and (4) as per the building code, the water in the damper can also be used for firefighting, which makes the damper more versatile. Based on the considerations above, a bidirectional TLSD is used for this project.

8.2.11.4.2 Damping type A bidirectional TLSD is designed on mechanical penthouse level on 67th floor of the building. To get the required damping, the damper is designed to have effective water mass ratio (equals to the ratio between effective water mass and the building modal mass) of 2.4% and 2.5% in two translational directions. The sloshing frequency in each direction is tuned to its optimal target frequency which is similar to the natural frequency of the building. To generate optimal damping of the damper, two 45% porous internal baffles are installed at approximately 0.4L and 0.6L locations along the length of the tank in both directions, where L indicates tank length. The performance of sloshing damper is verified through two series of 1:10 scale dynamic tests, each of which represents corresponding directional motion of the damper.

8.2.11.4.3 Structural and damping design Code and guidelines. The damper is designed to provide required damping under wind loads measured from 1:400 scale HFFB measurement of the building in the wind tunnel. The 10-year return period hourly wind pressure of 0.39 kPa was used for the test based on OBC (2006). Wind tunnel tests were conducted to meet or exceed the requirements of “Wind Tunnel Studies of Buildings and Structures,” ASCE Manual 7 Engineering Practice Note 67 (ASCE, 1999). Analysis modeling and software. Before the dynamic test, linear time history analyses of the sloshing damper-building coupled system were performed to predict the performance of the damper in both directions. Design principles. The damper is designed to be able to mitigate building displacements much less than 270 mm, which is 10-year return period peak displacement at the top occupied floor. The additional damping values added to the building are estimated to be approximately in 3.5% and 4.0% of critical for X and

8.2 Mass Damping Systems Case Studies

Y directions, respectively, which are to be added to the inherent structural damping of 2.0% of critical. The damper is tuned to the lowest lateral vibration mode in each direction. Design phase considerations. At the preliminary design stage, preliminary dimensions of the tanks and water height as well as the additional damping were estimated. To verify the additional damping, a dynamic test was performed in 1:10 scale under the dynamic wind loads measured in wind tunnel. The structural and architectural design was conducted to ensure the structural safety of supporting structures and walls. A long-term field monitoring of the building motion is being conducted and building accelerations will be verified through a full-scale measurement of building motions under strong wind events, and the as-built building frequencies will be measured for more accurate tuning. Documents for the damper design include detailed drawings of the damper and waterproofing specifications.

8.2.11.4.4 Architectural integration strategy The damper is placed on 67th mechanical penthouse floor to give enough space required for installation and maintenance of the damper.

8.2.11.4.5 MEP integration strategy Water supply and drainage piping are integrated into the tank to fill and drain the tank as necessary for the maintenance purposes. The tank is internal to the building so that no additional heating is required to maintain water from freezing. Water is treated with light concentration of chlorine to prevent major fouling.

8.2.11.4.6 Elevator and other mechanical devices integration strategy No information is available.

8.2.11.4.7 Experimental tests 1:10 scale dynamic model tests are performed to verify the damping effect of the damper.

8.2.11.4.8 Monitoring system Long-term full-scale building monitoring is being conducted before and after the installation of the tank, to verify the efficiency of the damper during strong winds.

8.2.11.4.9 Maintenance strategy Due to its passive nature, TSDs have lower maintenance requirements as compared to other systems. Principle maintenance tasks consist of two parts, keeping the proper water level and ensuring waterproofing during the lifetime of the damper. To maintain the water level, regular visual inspection through a transparent vertical glass or automatic water level controller with automatic water filling system has been used. Regular inspection is planned to avoid leaking. The tank

845

846

CHAPTER 8 Case studies of tall buildings

will be drained every 2 years for a direct inspection of the waterproofing liner. The screen components of the tank are corrosion-resistant stainless steel columns with structural plastic screens. The tank is refilled to the water level within a couple of days after the complete drainage.

8.2.11.4.10 Development process Based on wind tunnel testing and preliminary damper drawing, preliminary damper design was prepared by wind laboratory followed by structural design of the supporting structures and slabs. Detailed drawings prepared by the wind laboratory were reviewed by the design office before the installation of the damper.

8.2.11.5 Comparison of different design strategies 8.2.11.5.1 Structural options considered No options were considered.

8.2.11.5.2 Damping solution considered Different types of damping were considered including VE damper. However, due to availability of space, dual requirement for firefighting, and low cost, the sloshing damper solution was considered to be optimum.

8.2.11.5.3 Cost
benefit analysis The structural engineer and wind consultant went through a series of evaluations to determine what degree of stiffening the building and adding mass (i.e., changing the structural properties) would be required to reduce or eliminate the size of damper. However, the architectural layout of the building did not allow for significant alteration in the structural footprint. Therefore, the damper was relied upon to reduce the accelerations without significant change to the structure. Damping cost. Estimated to be 1 million $CDN (almost US$800,000) including internal components.

8.2.11.6 Lesson learned and recommendations 8.2.11.6.1 Difficulties in the design Building accelerations were determined only after the gross floor plate and structural layout was determined and the building was marketed. Better structural solutions and smaller damper solutions could have been available with earlier consideration of acceleration risks.

8.2.11.6.2 Design innovative solutions Use of 1:10 scale model testing and nonlinear analysis to derive most efficiency from liquid damper. Use of stainless steel columns and structural plastic to avoid corrosion issues over the life of the building. Use of high-quality liner to avoid leakage problems.

8.2 Mass Damping Systems Case Studies

8.2.11.6.3 Possible improvements Raise the tank by a small amount (30.5 cm) above the mechanical room slab to provide a secondary method to inspect for leakage.

8.2.12 ONE BLOOR STREET EAST, TORONTO, CANADA 8.2.12.1 Project data The major building data (Fig. 8.231) are summarized as follows: • •

• •

Year of completion: 2015 Developer/contractor Tuscany Ridge Developments (Canada) Great Gulf Group (Canada) Architectural designer: Hariri Pontarini Architects (Canada) Structural designer: Jablonsky, AST and Partners (Canada)

FIGURE 8.231 One Bloor Street East: building overview.

847

848

CHAPTER 8 Case studies of tall buildings

• • • • • • • •

Damping supplier: Gradient Wind Engineering Inc. (Canada) Testing laboratory: Gradient Wind Engineering Inc. (Canada) Height: 255 m Stories: 75 above grade Gross area of the tower: 68,634 m2 (738,770 ft2) Building function: Condominium residential/retail Structural material: Conventional reinforced concrete—Steel % not available Floor plan area: No information available

8.2.12.2 Introduction/history No information is available.

8.2.12.3 Structural system The structural system of the tower is composed of ductile concrete core shear wall with flat slabs (Figs. 8.232 and 8.233).

8.2.12.3.1 Building fundamental periods Natural vibration periods of the tower are calculated to be 7.0 in shorter direction in plan, 6.5 in longer direction, and 4.5 seconds in torsion based on the numerical structural model. Average mass density (per volume) of the building is approximately 300 kg/m3 which equals to weight density of 19 lb/ft3.

FIGURE 8.232 One Bloor Street East: typical floor plan.

8.2 Mass Damping Systems Case Studies

FIGURE 8.233 One Bloor Street East: typical elevation.

849

850

CHAPTER 8 Case studies of tall buildings

8.2.12.3.2 Damping strategy utilized Wind tunnel tests predicted that the building would experience high-amplitude vibration in both lateral directions. Additional damping of 2.3% was required to reduce the accelerations to the acceptable level based on industry guidelines. Since the accelerations in both translational directions were high, a bidirectional TLSD was designed at the top of the building above the mechanical floors to get an efficient damping within a given space.

8.2.12.3.3 Additional damping provided by the damping system Additional damping introduced by the TLSD was estimated to be approximately 2.5% in both lateral directions. The additional damping value of the damper was verified through a 1:10 scale dynamic building-damper testing in dynamic test rig facility at Gradient Wind Engineering Inc.

8.2.12.3.4 Building cost versus damping cost No information is available.

8.2.12.3.5 Building code Following is a list of the major standard design codes used in the project: • • • • • • •

CSA S16-09—design of steel structures (CSA, 2009) CSA 23.3-04—design of concrete structures (CSA, 2004) Ontario Building Code, 2006, 2012 (OBC, 2006, 2012) National Building Code of Canada 2005, 2010 (NBCC, 2005, 2010) ASCE 7-05 (ASCE, 2005) Eurocode 1 (ASCE, 2010) International Standard Organization (ISO, 2007)

8.2.12.3.6 Peer-reviewed Project None.

8.2.12.3.7 Design forces Wind loads were determined based on statistical model of the wind speed and direction data measured in Pearson International Airport in Toronto for the last 40 years. The measured wind speeds were calibrated to give a 50-year wind speed of 28.4 m/s at 10 m above grade corresponding to a dynamic pressure of 0.52 kPa for strength design and 50-year dynamic pressure of 0.48 kPa for deflection limit state design.

8.2.12.3.8 Expected performance The 1- and 10-year return period peak wind-induced accelerations at 75th floor, the top occupied floor, is predicted to be 14.1 and 25.5 mg, respectively. Peak torsional velocity is predicted to be 2.32 mrads/s. Based on the guidelines for building occupant comfort, the International Standards Organization (ISO, 2007) and

8.2 Mass Damping Systems Case Studies

CTBUH recommendation, the maximum 10-year return period resultant acceleration and torsional velocity are limited below 18 mg and 3 mrads/s, respectively. Based on the guidelines, the peak accelerations are not acceptable whereas torsional velocity is acceptable.

8.2.12.4 Damping overview To mitigate the acceleration levels below the serviceability limit of 18 mg, the building requires approximately 2.3% of additional damping in both Y and X directions in addition to the inherent structural damping.

8.2.12.4.1 Damping strategy Different types of artificial damping measures were investigated including TLSD and distributed damper. Following requirements were considered in damper design: (1) building interstory drift and corresponding deflection should be minimized, therefore the mechanism of the damper should be efficient with small excitation amplitude; (2) since the accelerations were high in both directions, damper should be activated in both directions; (3) the damper should be costefficient and easy to maintain. Based on the requirements above, a bidirectional TLSD was adopted for this project.

8.2.12.4.2 Damping type A bidirectional TLSD was designed atop the building to mitigate the building vibration to the acceptable level (Fig. 8.234). To get the required damping, the

FIGURE 8.234 One Bloor Street East: 3D scheme of bidirectional sloshing damper.

851

852

CHAPTER 8 Case studies of tall buildings

damper was designed to have effective water mass ratio (equal to the ratio between effective water mass and the building modal mass) of 2.5% and 2.7% in two translational directions. The sloshing frequency in each direction was tuned to its optimal target frequency which is similar to the natural frequency of the building. To generate optimal damping of the damper to dissipate the vibration energy more efficiently, two internal baffles with 40% porosity were installed at approximately 0.4L and 0.6L locations along the length of the tank in both directions, where L indicates tank length. The performance of sloshing damper was verified through two series of 1:10 scale dynamic tests corresponding to the principal directions of the damper.

8.2.12.4.3 Structural and damping design Code and guidelines. The damper was designed to provide required damping under the wind loads measured from 1:400 scale HFFB measurement of the building in wind tunnel. The 10-year return period hourly wind pressure of 0.36 kPa was used for the test based on NBCC (2005). Wind tunnel tests were conducted to meet or exceed the requirements of “Wind Tunnel Studies of Buildings and Structures,” ASCE Manual 7 Engineering Practice Note 67 (ASCE, 1999). Analysis modeling and software. Before the dynamic test, linear time history analyses of sloshing damper-building coupled system were performed to predict the performance of the damper in both directions. Design principles. The damper was designed to mitigate building displacements much less than 250 mm which is 10-year return period peak displacement at the top occupied floor. The damping added to the building was estimated to be approximately 2.5% of critical. The damper is tuned to the two lowest lateral vibration modes. Design phase considerations. At the preliminary design stage, preliminary dimensions of the tanks and water height as well as the additional damping were estimated. To verify the additional damping, a dynamic test was performed in 1:10 scale under the dynamic wind loads measured in wind tunnel. The structural and architectural design was conducted to ensure the structural safety of supporting structures and walls. Ongoing full-scale monitoring of building motion under strong winds will be used to determine as-built building properties and verify predicted accelerations and damper performance. Documents for the damper design include detailed drawings and waterproofing specifications.

8.2.12.4.4 Architectural integration strategy The damper is placed on top of the roof to give enough space required for the tank as well as to make best use of the upper floor space.

8.2.12.4.5 MEP integration strategy Water supply and drain plumbing are integrated to the tank to support regular maintenance. A minimal water heating is included to prevent water in the tank from freezing during the winter.

8.2 Mass Damping Systems Case Studies

8.2.12.4.6 Elevator and other mechanical devices integration strategy No information is available.

8.2.12.4.7 Experimental tests 1:10 small-scale dynamic model tests were performed to verify the damping effect of the damper.

8.2.12.4.8 Monitoring system Long-term full-scale field monitoring were performed starting immediately after the building was topped-off to determine building performance and damper efficiency pre- and postdamper commissioning. The monitoring system consists of a set of three accelerometers, data logger, and remote wireless communication system to monitor, record, and adjust measurement parameters as required over time.

8.2.12.4.9 Maintenance strategy Due to its passive nature, TSDs have lower maintenance requirements as compared to other systems. Principle maintenance tasks consist of two parts, keeping the proper water level, and ensuring waterproofing during the lifetime of the damper. To maintain the water level, regular visual inspection through a transparent vertical glass or automatic water level controller with automatic water filling system has been used. Regular inspection is planned to avoid leaking. The tank will be drained every 2 years for a direct inspection of the waterproofing liner. The screen components of the tank are corrosion-resistant stainless steel columns with structural plastic screens. The tank is refilled to the water level within a couple of days after the complete drainage.

8.2.12.4.10 Development process Based on wind tunnel testing and preliminary damper drawing, preliminary sloshing damper design was prepared by wind laboratory followed by structural design of the supporting structures and slabs. Detailed drawings prepared by the wind laboratory were reviewed by the design office before the installation of the damper.

8.2.12.5 Comparison of different design strategies 8.2.12.5.1 Structural options considered No significantly different structural options were considered but only refinements of the same system described previously, such as use of transfer slab thicknesses, outrigger system at rooftop.

8.2.12.5.2 Damping solution considered Different types of damping were considered including VE damper. However, in terms of cost-efficiency and maintenance aspects, TLSD was considered to be the best solution.

853

854

CHAPTER 8 Case studies of tall buildings

8.2.12.5.3 Cost
benefit analysis Damping cost. Estimated damper costs 1.2 million $CDN (almost US$960,000).

8.2.12.6 Lesson learned and recommendations 8.2.12.6.1 Difficulties in the design The large size of the damper tank and its integration with the window washing equipment was a challenge for the trades. As usual, the height of the building and the objective to obtain maximum salable floor space limited the vertical space to consider other damper options.

8.2.12.6.2 Design innovative solutions The damper tank used a novel method of waterproofing which is using BASF 300D—Masterseal 345 product, which is a spray-in-place application.

8.2.12.6.3 Possible improvements None available with the given constraints.

8.2.13 1151 WEST GEORGIA, VANCOUVER, CANADA 8.2.13.1 Project data The major building data (Fig. 8.235) are summarized as follows: • • • • • • • • • • • • •

• •

Year of completion: 2016 Developer/contractor: Holborn Group of Companies (Canada) Architectural designer: Arthur Erickson Architectural Corporation, Musson Cattel Mackey Architects, DYS Architecture (Canada) Structural designer: John Bryson & Partners (Canada) Damping supplier: Gradient Wind Engineering Inc. (Canada) Testing laboratory: Gradient Wind Engineering Inc. (Canada) Height: 187 m Interstory height: 2.9
3.0 m Slenderness ratio: 1:6.2 Stories: 60 above grade Gross area of the tower: Approximately 54,000 m2 Building function: Condominium residential/hotel Structural material Concrete classes: Wall concrete strength 35 MPa Structural steel: Information NA Dampers: Two-tank TLSD on rooftop (level 70) Rebar: Information NA Structural system: ductile concrete core shear wall with flat slabs Floor plan area: approximately 669 m2

8.2 Mass Damping Systems Case Studies

FIGURE 8.235 1151 West Georgia: building overview.

8.2.13.2 Introduction/history A supplemental damping system has been investigated later in the design stage of the building due to late design changes. The most compact method for late implementation of the required supplemental damping system has been investigated. The damping systems include the TLCD, single-tank TLSD, and multiple (two or four)-tank TLSD. The two-tank TLSD was selected for the building based on a review of efficiency, constructability, and maintenance.

8.2.13.3 Structural system The structural system is composed of a ductile concrete core shear wall with flat slabs.

8.2.13.3.1 Building fundamental periods The natural vibration periods of the tower are 7.0, 6.5, and 2.3 seconds in two translations and rotation, respectively. Average mass density (per volume) of the building is approximately 300 kg/m3, which corresponds to weight density of 19 lb/ft3 (Fig. 8.236).

8.2.13.3.2 Damping strategy utilized Wind tunnel tests predicted (Fig. 8.237) that the building will experience highamplitude vibration especially in NS direction. Additional damping of 2.0% was required to bring the accelerations down to the acceptable level based on industry guidelines. Two identical 1D TLSDs were designed atop the building.

855

CHAPTER 8 Case studies of tall buildings

Mode 1

Mode 2

Mode 3

700

700

600

600

600

500

500

500

400

400

400

300

200

Elevation (ft)

700

Elevation (ft)

Elevation (ft)

856

300

200

200 x y

x y

Rotation (z)

Rotation (z)

x y 100

–0.4 –0.2

0

0

300

Rotation (z)

0.2 0.4 0.6 0.8 Mode shape

Mode 1:T1 = 7.0 s

1

1.2

100

–0.2

0

0

0.2 0.4 0.6 Mode shape

0.8

1

1.2

100

–0.6 –0.5 –0.4 –0.3 –0.2 –0.1 Mode shape

Mode 2:T2 = 6.5 s

0

0

0.1

Mode 3:T3 = 2.3 s

FIGURE 8.236 1151 West Georgia: fundamental frequencies and mode shapes.

FIGURE 8.237 1151 West Georgia: wind tunnel test model (looking downstream).

8.2.13.3.3 Additional damping provided by the damping system The additional damping introduced by the TLSDs was approximately 2.0% in dominant NS direction. The additional damping value of the damper has been verified through a 1:10 scale dynamic building-damper testing in dynamic test rig facility (Fig. 8.242).

8.2.13.3.4 Building cost versus damping cost Cost of damper estimated to be 0.6 million $CDN (almost US$480,000). The building cost is unavailable.

8.2 Mass Damping Systems Case Studies

8.2.13.3.5 Building code Following is a list of the major standard design codes used in the project: • • • • • •

CSA S16-09—design of steel structures (CSA, 2009) CSA 23.3-04—design of concrete structures (CSA, 2004) National Building Code of Canada 2010 (NBCC, 2010) ASCE 7-05 (ASCE, 2005) Eurocode 1 (CEN, 2010) ISO: ISO 6897 (ISO, 1984), ISO 10137 (ISO, 2007)

8.2.13.3.6 Peer-reviewed project None.

8.2.13.3.7 Design forces Wind loads were determined based on statistical model of the wind speed and direction data measured at Vancouver International Airport for last 40 years. The measured wind speeds were calibrated to give a 50-year return period wind speed of 28.8 m/s at 10 m above grade corresponding to a dynamic pressure of 0.48 kPa, and 10-year dynamic pressure of 0.36 kPa based on National Building Code of Canada (NBCC, 2010). The building was designed for both wind and seismic. Wind controlled the design of some structural shear walls and building serviceability limits for deflections and accelerations.

8.2.13.3.8 Expected performance The 1- and 10-year return period peak wind-induced accelerations at the top occupied floor are predicted to be 12.9 and 22.1 mg, respectively, without supplemental damping. The largest part of the acceleration is contributed by Y direction (NS directional) component. Torsional velocity remained small due to the high torsional frequency of the building and compact shape. Based on the target accelerations guidelines of 18 mg for the 10-year acceleration, the building performance was not acceptable.

8.2.13.4 Damping overview To mitigate the acceleration levels below the serviceability limit of 18 mg based on a 10-year return period, the building requires approximately 2.0% of additional damping especially in Y direction in addition to the inherent structural damping.

8.2.13.4.1 Damping strategy The project had very limited space on the top floors due to the penthouse amenity spaces on the roof. Therefore, many different types of artificial damping measures were investigated including TLCD and single or multiple TLSDs to minimize the space for the damper. Based on the fact that the building motion is governed by one direction, the damper was able to be divided into multiple tanks to make most efficient use of the available spaces on the rooftop.

857

858

CHAPTER 8 Case studies of tall buildings

FIGURE 8.238 1151 West Georgia: preliminary sizing of TLCD options.

Fig. 8.238 illustrates a preliminary size of the considered TLCD system, which can provide 2.0% additional damping similar to that of TLSD. Due mainly to the constructability and maintenance, these TLCD options were not pursued further.

8.2.13.4.2 Damping type A 1D TLSD with two identical tanks was designed for the building’s roof to mitigate excessive building vibration in the NS direction to an acceptable level. The damper was required to provide at least 2.0% additional damping to the structure. The damper was designed with an effective water mass ratio (equal to the ratio between effective water mass and the building modal mass) of approximately 2.0%. The sloshing frequency in each direction was tuned to its optimal target frequency, which is similar to the natural frequency of the building. To generate optimal damping force of the damper and dissipate the vibration energy more efficiently under the building vibration amplitude, two 55% porous internal baffles

8.2 Mass Damping Systems Case Studies

were installed at approximately 0.4L and 0.6L locations along the length of the tank, where L indicates tank length. The performance of the sloshing damper is verified from 1:10 scale dynamic tests of the building/sloshing damper coupled system.

8.2.13.4.3 Structural and damping design Code and guidelines. The damper is designed to provide required damping under the wind loads measured from a 1:400 scale HFFB model of the building in wind tunnel. The 10-year return period hourly wind pressure of 0.36 kPa is used for the test based on NBCC (2005). The wind tunnel test is conducted to meet or exceed the requirements of “Wind Tunnel Studies of Buildings and Structures,” ASCE Manual 7 Engineering Practice Note 67 (ASCE, 1999). Analysis modeling and software. Before the dynamic test, linear time history analyses of the sloshing damper-building coupled system were performed to estimate the performance of the damper before the dynamic test. Design parameters such as spring stiffness and mass as well as the time history of actuator forces were determined and verified from the analysis. In the analysis, a scaled building model and a damper were modeled as a two degree-of-freedom system in model scale. Wind loads measured from a wind tunnel test were scaled in magnitude and time, and applied to the model for the time-domain analysis. In the analysis, Gradient Wind’s in-house Matlab program was used. In the program, TLSD was modeled as a single degree-of-freedom equivalent secondary mass, of which equivalent mass and screen damping forces were estimated based on the analytical solution of potential flow (Tait, 2008). Design principles. The building displacements to be mitigated are much lower than 200 mm, which is the highest measurable displacement that the building would experience on average at the top occupied floor in 10 years. The additional damping added to the building is estimated to be approximately 2.0% of critical, which is to be added to the assumed inherent structural damping of 1.75% of critical. The damper is optimally tuned to the NS directional building vibration with a natural vibration frequency of 0.153 Hz to create the largest damping effect. Based on the required supplemental damping in the direction of interest (Y direction), design parameters such as (effective) water mass ratios, water sloshing frequencies, and screen damping were evaluated to achieve the optimal damping. A water height was assumed and length of the tank was determined to obtain the target sloshing frequency based on the potential flow theory. Tank width was determined to achieve the required mass ratio. Since the length in one direction corresponds to the width in the other direction in bidirectional tanks, the two dimensions were determined iteratively to satisfy both the frequency and mass ratio by adjusting the water height. Design phase considerations. At the preliminary design stage, preliminary dimensions of the tanks and water height as well as the additional damping were estimated and provided to the architects and structural engineers. To verify the additional damping, a dynamic test was performed in 1:10 scale using a time

859

860

CHAPTER 8 Case studies of tall buildings

series of dynamic wind loads measured in the wind tunnel. A preliminary set of damper drawings was created based on the design parameters such as the location of the solid and porous screens and porosity of the screens. The structural and architectural design has been conducted to ensure the structural safety of supporting structures and walls. At the client’s request, short-term field monitoring was performed to measure the as-built building frequencies. Damper drawings were finalized based on the revised designed parameters and field measurement results. Documents for the damper design include detailed drawings of the damper and waterproofing specifications.

8.2.13.4.4 Architectural integration strategy To minimize the space of the damper, the damper was separated in two tanks and located atop the roof.

8.2.13.4.5 MEP integration strategy Information is not available.

8.2.13.4.6 Elevator and other mechanical devices integration strategy Information is not available.

8.2.13.4.7 Experimental tests The performance of sloshing damper was verified through testing of a 1:10 scale damper. Fig. 8.239 shows the dynamic test rig and 1:10 scale sloshing water tank

FIGURE 8.239 1151 West Georgia: overall view of pendulum-type dynamic test rig with a scaled TLD tank.

8.2 Mass Damping Systems Case Studies

used for the verification test. The pendulum mass and water tank represent approximately 25% of 1:10 scale building modal mass and TLD, respectively. The mass was designed to move freely in pendulum motion and was connected to springs representing building stiffness. The system was instrumented with accelerometers and load cells to measure the acceleration of the mass, the actuator excitation force, and the sloshing water force on the screens and tank end walls. To simulate the wind excitation force, wind load time history measured from the HFFB test was used. From the damper test, sloshing water frequency, screen damping, as well as supplemental damping of the damper, were measured. For the water depth-tolength ratio, theoretical sloshing frequencies based on potential flow theory matched very well with the test results. To determine the optimal porosity of internal screens that can provide the highest damping, tests were performed for a range of screen porosities. Lastly, supplemental damping introduced by the damper was measured, to be compared to the required damping.

8.2.13.4.8 Monitoring system Short-term (2 weeks) site monitoring was performed to measure the as-built building frequency. Two accelerometers were installed at the top mechanical floor to take measurements. The building’s natural frequencies in two orthogonal directions were measured by analyzing building motion during ambient vibration (i.e., by the operation of the elevator or mild winds). The measured building frequency was used to finalize the location of solid screens to optimally tune the damper to the building.

8.2.13.4.9 Maintenance strategy Due to its passive nature, TSDs have lower maintenance requirements as compared to other systems. Principal maintenance tasks consist of two parts: keeping the proper water level and ensuring waterproofing during the lifetime of the damper. To maintain the water level, regular visual inspection through a transparent vertical glass or automatic water level controller with an automatic water filling system was performed. Regular inspection is planned to avoid leaking. The tank will be drained every 2 years for a direct inspection of the waterproofing liner. The screen components of the tank are corrosion-resistant stainless steel columns with structural plastic screens. The tank is refilled to the water level within a couple of days after the complete drainage.

8.2.13.4.10 Development process Based on structural data and building geometry, respectively, provided by the structural engineers and the architects, wind tunnel testing of building motions and wind loads were performed. Since the building motions were predicted to be excessive, artificial damping strategies including TLCD and multiple TSD were investigated by the wind laboratory in coordination with the developer and the architects. The TLSD was selected as the optimum solution, requiring minimum

861

862

CHAPTER 8 Case studies of tall buildings

space and low maintenance costs. The structural design of the damper was performed by the structural engineers. Based on short-term monitoring immediately after topping off the building, the damper design was finalized and ready for the construction.

8.2.13.5 Comparison of different design strategies 8.2.13.5.1 Structural options considered Information is not available.

8.2.13.5.2 Damping solution considered Different types of damping were considered including TLCD and multiple (4 or 6)-tank TLSD. Application of TLCD was challenging, especially in waterproofing the column walls. When large water mass is required, the cross section of the column must be large, which reduces effective column length and makes water column behavior unclear at the bottom corners. Size-wise, there was little benefit of using TLCD over TLD. TLD with multiple (4 or 6) small tanks of various sizes were investigated; however, considering the small contribution to damping, small tanks are expected to create issues such as water leakage and maintenance and this option was not pursued further. Considering space limitation and the aspect of maintenance, a TLSD with two tanks was considered to be the best solution for the building.

8.2.13.5.3 Cost
benefit analysis The structural engineer and wind consultant went through a series of evaluations to determine what degree of stiffening the building and adding mass (i.e., changing the structural properties) would be required to reduce or eliminate the size of damper. However, the architectural layout of the building did not allow for significant alteration in the structural footprint. Therefore, the damper was relied upon to reduce the accelerations without significant change to the structure. Damping cost. Damper cost estimated to be 0.6 million $CDN (almost US$480,000).

8.2.13.6 Lesson learned and recommendations 8.2.13.6.1 Difficulties in the design Design difficulties centered around very tight spaces and headroom available for the damper due to late architectural design changes. Wind/damper consultant went through many damper design iterations to select a damper of constrained size to achieve requirement for damping.

8.2 Mass Damping Systems Case Studies

8.2.13.6.2 Design innovative solutions These included the exhaustive damper screen testing and analysis to fit the damper into the allowable spaces.

8.2.13.6.3 Possible improvements Develop a new damper concept that would be smaller and cheaper than any of the currently available options including TMD, pendulum, and double pendulum dampers.

8.2.14 SHANGHAI TOWER, SHANGHAI, CHINA 8.2.14.1 Project data The major building data (Fig. 8.240) are summarized as follows: • • •

•

•

• • • • • • • •

• •

Year of completion: 2016 Developer/contractor: Shanghai Tower Construction & Development Co., Ltd (China) Architectural designer: Gensler (design) (United States) Architectural Design & Research Institute of Tongji University (Group) Co., Ltd (local design institute) (China) Structural designer: Thornton Tomasetti (design) (United States) Architectural Design & Research Institute of Tongji University (Group) Co., Ltd (local design institute) (China) Damping supplier RWDI (design) (Canada) SRIM (China) Testing laboratory: Wind tunnel tests performed at RWDI Inc. Height: 632 m Interstory height: 4.5 m for office floors and 4.3 m for hotel floors Slenderness ratio: 1:7.6 Stories: 128 above grade; 5 below grade Gross area of the tower: 574,058 m2 Building function: Hotel, office Structural material Composite concrete core Steel mega-frame (mega-columns with box belt trusses) Steel outrigger trusses Damping type Pendulum TMD with an eddy current damping system Floor plan area: Approximately 6600 m2 at the bottom

863

864

CHAPTER 8 Case studies of tall buildings

FIGURE 8.240 Shanghai Tower: building overview.

8.2.14.2 Introduction/history Located in lot Z3-1 and adjacent to Jin Mao Tower and Shanghai World Financial Center, Shanghai Tower rises 632 m above the ground as a landmark on the city skyline. The Shanghai Tower will house mixed use of Class A office space, entertainment venues, retail stores, a conference center, a luxury hotel, and cultural amenities. The 5-story deep basement serves retail, MEP, and parking spaces. Shanghai Tower has a unique twisting skin, but inside it takes the form of nine cylindrical buildings stacked one atop another, including a business zone, five office zones, two hotel/apartment zones, and one top zone with sightseeing or observation floors. Each zone can be considered as an independent city or village with communal space at an amenity level where the slab extends to reach

8.2 Mass Damping Systems Case Studies

the outer twisting fac¸ade. The regular tower floor plate at each zone is in circular shape with diameter varies from 82.2 m at bottom to 46.5 m at top. The stackedzone tower concept within an exterior fac¸ade tapering and twisting with height creates a spectacular architectural design.

8.2.14.3 Structural system Shanghai Tower adopts the core outriggers mega-frame lateral force resisting system, which consists of three parts: composite concrete core, exterior mega-frame (composite super-columns with steel box belt trusses), and steel outrigger trusses (Figs. 8.241
8.243). The gravity system uses steel floor beams and perimeter gravity columns to pick up the floor gravity loads from the composite metal decks, and transfer the loads into the super-columns through the belt trusses (Figs. 8.241
8.243).

FIGURE 8.241 Shanghai Tower: typical structural framing plans.

865

866

CHAPTER 8 Case studies of tall buildings

FIGURE 8.242 Shanghai Tower: typical MEP/refugee structural framing.

8.2.14.3.1 Building fundamental periods The building fundamental vibration modes are shown in Fig. 8.244. The first six mode periods are the followings: • • • • • •

9.04 seconds 8.97 seconds 5.79 seconds 3.34 seconds 3.21 seconds 2.70 seconds

first mode is mainly flexural along Y direction second mode is mainly flexural along X direction third mode is mainly torsional forth mode fifth mode sixth mode

8.2.14.3.2 Damping strategy utilized Either increasing the tower stiffness or mass is far less effective than a supplemental damping system. Therefore, a 1000-ton simple pendulum TMD with an

8.2 Mass Damping Systems Case Studies

FIGURE 8.243 Shanghai Tower: typical elevation.

867

868

CHAPTER 8 Case studies of tall buildings

FIGURE 8.244 Shanghai Tower: fundamental frequencies and mode shapes.

eddy current damping system and a viscous damping snubber system was utilized for increasing the damping in the structure.

8.2.14.3.3 Additional damping provided by the damping system TMD equivalent damping ratio for the first two modes is equal to 2.3%.

8.2.14.3.4 Building cost versus damping cost Damping cost not to be disclosed by the owner.

8.2.14.3.5 Building code China building codes.

8.2.14.3.6 Peer-reviewed project Structural design of the tower was reviewed and approved by the China National Expert Review Panel.

8.2.14.3.7 Design forces Seismic zone: degree 7; 50-year reference wind pressure: 0.55 kPa.

8.2.14.3.8 Expected performance The TMD is designed to bring down 45% peak acceleration induced by lower return period wind. For 10-year return period wind, the acceleration of top occupied floor is reduced from 19 to 11 mg which meets the requirement of a residential tower in China building code. For 1-year return period wind, the acceleration of top occupied floor is reduced from 5 to 2.8 mg which achieves the H10 comfort level.

8.2 Mass Damping Systems Case Studies

8.2.14.4 Damping overview The lateral resisting system consisting composite concrete core, mega-frame (mega columns with box belt trusses), and outrigger trusses makes Shanghai Tower effectively against wind load and seismic load. With a twisting and tapering aerodynamic shape that greatly reduces wind load, the structural design governed by story drift under 50-year wind load and vibration acceleration under 10-year wind load have been optimized from the results of wind tunnel tests done by RWDI. To forge a high-end hotel on the upper levels, the design team would like to make this tower meet the requirement of the most stringent comfort criteria in the world, the H10 level of Architectural Institute of Japan (AIJ), which means only about 10% people could sense the vibration (Fig. 8.245). So it is determined to implement a damping system to achieve a cost-effective design. At the stage of concept design, RWDI proposed a 1200-ton double pendulum TMD which requires less height than a simple pendulum style. By coordinating with the design team, it is determined to use the simple pendulum style as it is more concise and can highlight the sculpture stands on TMD mass box (Figs. 8.246 and 8.247). After SRIM introduced the eddy current damping system into the design, RWDI studied and developed new tools for design, analysis of this new damping system. Substituting the primary viscous damping devises by eddy current damping system, the TMD mass was optimized from 1200 to 1000 tons. The final design TMD performance achieves 45% reduction in acceleration with eddy 35 With TMD

Total peak acceleration (mg)

30 Without TMD

25 RWDI office 20 RWDI residential

ISO office

15

ISO residential

10 5

Typical time between occurrences

FIGURE 8.245 Shanghai Tower: acceleration versus return period.

Ye ar s 50

Ye ar s 30

Ye ar s

15

Ye ar s 10

s

s

Ye ar 5

s

Ye ar

3

Ye ar 2

on th s 12 ont h M s on th s

M

M

9

6

on th M s on th s

M

3

2

1

M

on th

0

869

870

CHAPTER 8 Case studies of tall buildings

FIGURE 8.246 Shanghai Tower: TMD general scheme.

current damping system (Table 8.41). See Fig. 8.250 for acceleration versus return period.

8.2.14.4.1 Damping strategy Distributed damping systems have been widely used in seismic applications for mid-rise buildings (15
30 stories). Service level wind load (i.e., 1- and 10-year wind events) is usually much smaller than seismic load. This leads to limited levels of interstory drift/deformations and makes the implementation and optimization of distributed damping systems a challenging exercise. To extract energy from the limited levels of wind-induced relative deformations between floors or between the core and the outer gravity columns at service level winds the designer needs to install a considerable number of distributed damper units along the height of the building at locations where you can get the most out of them. TMD is a supplemental inertia system that absorbs vibration energy from the building and amplifies its own vibration. It is able to provide effective control force under small deformation of the building. Therefore to mitigate wind-induced acceleration, TMD performs better than distributed damping systems.

8.2 Mass Damping Systems Case Studies

FIGURE 8.247 Shanghai Tower: TMD view.

Table 8.41 Shanghai Tower: Peak Acceleration Reduction With and Without TMD Peak Acceleration (mg) Total
[X, Y, and Torsional Component]

Reduction Factor

Return Period (Years)

Without TMD

With TMD

(1-With/Without TMD)

1 5 10

5.0
[4.4, 3.7, 0.8] 12.0
[9.8, 9.1, 0.8] 19.0
[15.0, 15.0, 2.1]

2.8
[2.4, 2.1, 0.8] 6.5
[5.4, 5.0, 1.5] 11.0
[8.4, 8.1, 2.1]

45% 45% 44%

8.2.14.4.2 Damping type Primary damping system. An eddy current damping system was firstly utilized in a large TMD in the world. When a magnetic field moves through a conductor the movement induces an eddy current in the conductor. The flow of electrons in the conductor immediately creates an opposing magnetic field which results in

871

872

CHAPTER 8 Case studies of tall buildings

damping of the magnet and produces heat inside the conductor similar to heat buildup inside of a power cord during use. When copper plate (conductor) thickness, magnet property, and area are determined, the eddy current damping force is determined by the gap between magnet and copper plate. The relationship of force and velocity is linear (e.g., Force 5 C 3 V) when a constant gap is maintained. The linear damping can provide consistent performance in a wide range of return period wind events. However, for this project, the TMD amplitude in high return period wind or seismic events will be over 2 m and become unacceptable if the optimal performance-based linear damping is used. In order to control the large TMD displacement, an increasing damping ratio with TMD amplitude is needed. This can be achieved by adjusting copper plate thickness and the gap between magnet and copper plate, which will generate a TMD-position-dependent nonlinear damping. The hysteresis curve at different amplitudes changes from ellipse shape (small amplitude) to bow tie shape (large amplitude) (see Fig. 8.248). This unique feature allows TMD to achieve optimal performance under low return period wind event while its amplitude can be controlled under extreme event with snubber system involved. Snubber system. A snubber system consisting of a steel ring, four supporting steel beams, and eight VDDs is designed to begin to work at a certain amplitude to absorb energy under extreme event.

FIGURE 8.248 Shanghai Tower: TMD eddy current damping system hysteretic curve.

8.2 Mass Damping Systems Case Studies

8.2.14.4.3 Structural and damping design The lateral force resisting system consists of the composite concrete core, composite exterior mega-frame, and the steel outrigger trusses (Fig. 8.243). The core wall resists most of lateral shear and serves as the fundamental defense line to prevent building collapse under severe seismic events. There are eight super-columns up to Zone 8 and four diagonal columns up to Zone 5 which are reinforced with built-up steel plates of approximately 4%
6% of column cross area. All of them work together with eight sets of 2-story-high double belt trusses to form the “exterior mega-frame” which serves as a second line of defense required by China code (Fig. 8.241). Six sets of 2-story-high steel outrigger trusses are placed at the MEP floors (Fig. 8.247). When connected to the core through outriggers, the super-columns provide large bending stiffness for the tower structure. The outrigger trusses and belt trusses help the structural system to be stiff enough to meet the stringent story drift ratio limit required by China code. The maximum story drift is h/505 under 50-year wind load and h/623 under frequent seismic load (a 50-year seismic event). One-story-high radial trusses cantilever at the upper MEP level to support slab areas beyond the super-columns. Those radial trusses also support the exterior fac¸ade system. Code and guidelines. The design follows the China building codes (GB50011, 2010). International building codes are also referred to whenever appropriate. Analysis modeling and software. The ETABS and SAP software developed by Computer and Structures, Inc. (CSI, 2016) is used to build the structural analysis model. Thornton Tomasetti, Inc. developed in-house software and spreadsheets to perform the structural design based on the analysis results. RWDI used in-house programs for the TMD analysis and the SolidWorks for the TMD design. Design principles. No information is available. Design phase considerations. No information is available.

8.2.14.4.4 Architectural integration strategy The TMD of Shanghai Tower is located at the center of the tower. The TMD design and location is coordinated with the architect to avoid interference with the functional area. In addition, a sculpture has been designed and integrated with the TMD so that the TMD becomes a decoration and architectural feature of the building top for the visiting public audiences.

8.2.14.4.5 MEP integration strategy The Shanghai Tower’s TMD is located within the center of the crown structure above most of the tower’s mechanical equipment. But some MEP equipment (e.g., cooling towers and firefighting water tanks) are hidden behind the TMD space.

873

874

CHAPTER 8 Case studies of tall buildings

8.2.14.4.6 Elevator and other mechanical devices integration strategy The TMD design needs to be coordinated with the elevator overrun at the top of the building.

8.2.14.4.7 Experimental tests Large-scale model of the tower was tested on the shake table to evaluate the performance of the tower structure during earthquake. The tests proved that the tower structure designed by Thornton Tomasetti, Inc., is able to achieve the targeted performance goal during strong earthquakes. A large-scale model of the TMD was also designed and built to verify the effectiveness of the novel eddy current dissipation system.

8.2.14.4.8 Monitoring system A comprehensive monitoring program has been implemented with sensors on the tower to monitor the wind and seismic loads, as well as the building performance during the strong wind and earthquake events. During the Tuning and Commissioning phase of the TMD installation, additional accelerometers were temporarily employed on the building and TMD mass itself to collect data from the forced excitation tests.

8.2.14.4.9 Maintenance strategy The TMD is designed to be largely maintenance free but periodic inspection is recommended to insure that the TMD is in good working order. If a strong wind or seismic event occurs, the TMD should be inspected to verify all components are still performing satisfactorily.

8.2.14.5 Comparison of different design strategies 8.2.14.5.1 Structural options considered The success of a super-tall building design relies on the selection of an efficient lateral load resisting system. The cost of lateral system accounts for a large portion of construction cost, so several structural options, including tube-in-tube, mega-frames, a few hybrid systems, and core outrigger mega-frame scheme, were developed at the beginning of the project for the owner, the architect, and the cost estimator to evaluate. Outrigger mega-frame scheme was determined to be the optimal system for the Shanghai Tower. After the structural option was selected, engineers made continuous effort to optimize the structure through all phases of the project. The goal was to find the most economical system using the least quantity of building material without compromising the architectural functions. At the beginning of the design stage, outriggers were placed at every MEP floor in order to maximize the structural lateral stiffness to meet the stringent story drift requirement in China code. The outrigger locations along building height were extensively studied and optimized. Engineers found the outriggers at low zones are effective in reducing

8.2 Mass Damping Systems Case Studies

the building fundamental period, while upper outriggers contribute more to control the story drifts at upper zones. The first and third outriggers were removed and achieved a saving of 3500 tons of steel.

8.2.14.5.2 Damping solution considered Multiple iterations of the TMD were considered and revised based upon changing objectives of the architect and owner. Initially, there was a premium placed on vertical space occupied by the TMD, and a more compact arrangement was designed that utilized an intermediate frame to essentially cut the required height by a factor of 2. As the vision to create an esthetically pleasing TMD matured, this design was superseded by a simpler and more elegant single pendulum, with very tall cables that can now be seen to stretch up through the sculpted ceiling. Further work by the architect resulted in the addition of the sculpture to the top of the otherwise unremarkable (but large) ballast mass portion of the TMD. Ownership was interested to showcase advanced technology and to distinguish their building and TMD from others by using the world’s largest eddy current damping system instead of the conventional sealed hydraulic damping cylinders. This results in extremely low friction, such that the TMD can be seen operating more of the time, into wind speeds of even lower speed. This increased the cost, but was approved to improve the attractiveness and interest from the visiting public audience.

8.2.14.5.3 Cost
benefit analysis No information is available.

8.2.14.6 Lesson learned and recommendations 8.2.14.6.1 Difficulties in the design The city of Shanghai is on the east coast of China and in a typhoon-prone zone. The high design wind load combined with the stringent stiffness requirements of the China code makes the lateral force resisting system design very challenging. Like other super-tall buildings, the foundation mat design of Shanghai Tower is a big challenge due to large gravity force and large overturning forces from wind and seismic loads. Soft soil conditions at site reinforced that challenge. With nine layers of sands and soft clays alternating to at least 120 m below grade, bedrock is considered beyond reach for practical construction purposes. Because the top 15 m is very soft silty clay, the site for seismic design is considered as Type IV—the most unfavorable class according to the China code and roughly comparable to Site Class “F” IBC, 2012.

8.2.14.6.2 Design innovative solutions When the Shanghai Tower was being designed, it is by far the tallest tower in China. The local engineers, contracts, and authorities have scarce experience with

875

876

CHAPTER 8 Case studies of tall buildings

buildings of that height. Thornton Tomasetti, Inc. adopted a lot of new methods and technologies for the first time in China: •

•

•

•

An innovative “core outrigger mega-frame” lateral system is adopted to meet conservative China code requirements for lateral stiffness and member capacities to realize the iconic architectural profile effectively. The 6-m thick CIP pile supported mat foundation, enhanced by the concrete fin walls at basement levels, help distributes loads more evenly to the soft supporting soil underneath. End grouting is provided at tip of piles to increase pile capacity and reduce settlement. Optimization of the tower lateral system is a continuous process through all design phases: 20% of wind load reduction is achieved through fine-tuning of tower profile; 13,000 tons of steel saving is achieved through extensive optimization of outrigger trusses and structural steels encased in super-columns. Advanced performance-based design is used to verify tower performance under different seismic hazard levels through nonlinear dynamic time history analysis.

8.2.14.6.3 Possible improvements Many aspects of the TMD were done to be unique in the world, and are unlikely to be repeated when cost is the chief concern. It is a one-of-a-kind design and many new features were developed especially for this project. Coordination between many participants was required throughout, many of whom had never seen a damping system before. This education and coordination required great vigilance, particularly in light of the mixed Chinese and English language capabilities of all parties. This coordination was ultimately successful, but would be expected to go even more smoothly in a repeated exercise now that the first-time experience has been completed.

8.2.15 THE INDEPENDENT, AUSTIN, TEXAS, UNITED STATES 8.2.15.1 Project data The major building data (Fig. 8.249) are summarized as follows: • • • • • • • • • • • •

Year of completion: 2018 Estimated Developer: Aspen Heights Partners/CIM Group (United States) Contractor: Balfour Beatty (United States) Architectural designer: Rhode Partners (United States) Structural designer: LAM 1 DCI (United States) Damping supplier: Fiber Technology Corporation (United States) Testing laboratory: Windtech Consultants (Australia) Height: 209 m Interstory height: 3.28 m Stories: 62 above grade Gross area of the tower: 88,258 m2 Building function: Residential

8.2 Mass Damping Systems Case Studies

FIGURE 8.249 The Independent: building overview.

• •

Structural material: Reinforced concrete Floor plan area: approximately 1256 m2 per floor

8.2.15.2 Introduction/history The luxury condominium tower, developed by Aspen Heights Partners, is set to become the tallest all concrete residential tower west of Mississippi River (Fig. 8.250).

8.2.15.3 Structural system Flat plate posttensioned slabs with single ordinary RC shear wall core for tower and additional shear walls for lower podium levels. Tower floor plate is supported by the core, eight mega columns and four corner columns with varying cantilevered slabs. Use of rooftop outrigger and viscous style liquid damper tank (Fig. 8.251).

877

FIGURE 8.250 The Independent: elevation diagram and plan drawings.

8.2 Mass Damping Systems Case Studies

FIGURE 8.251 The Independent: elevation and section drawing illustrating the structural system.

The structural system utilizes the results of advanced wind tunnel tests. The structural design solution results in resident comfort during high wind events. The Independent will integrate: an efficient core (aspect ratio approximately 17), rooftop outrigger that engages mega columns, damper tank technology, and highstrength reinforced steel link beams. Other innovative solutions include high-strength rebar, extreme cantilever amenity level, extended cantilever floors with tension rods, and cantilevered pool.

8.2.15.3.1 Building fundamental periods The estimated periods for the first three modes of vibration: 6.6, 6.3, 2.1 seconds. The natural frequencies will be measured and the finite element analysis model calibrated once the structure has been completed and prior to finalizing the dimensions of the liquid damper. The corresponding mode shapes are shown in Fig. 8.252.

8.2.15.3.2 Damping strategy utilized Initially set out to utilize a proposed fire hydrant tank as a sloshing damper to reduce accelerations, which were estimated to just fall within the occupant

879

880

CHAPTER 8 Case studies of tall buildings

FIGURE 8.252 The Independent: fundamental frequencies and mode shapes.

comfort criterion for residential buildings based on ISO 10137 (ISO, 2007) and a 1% inherent structural damping assumption for the estimated 1-year return amplitude of sway. The design was later modified to enable this passive liquid damper design to perform over a broad range of frequencies with less sensitivity to tuning (the idea arrived at by Windtech Consultants from concept of a dashpot). This resulted in as much as 45% efficiency with a tuning ratio as low as 0.8. A 25% efficiency was assumed when determining the contribution to the ULS design loads.

8.2.15.3.3 Additional damping provided by the damping system The liquid damper, when tuned, can provide up to 2% additional damping to the structure but can also provide 0.9% damping at a tuning ratio of 0.8.

8.2.15.3.4 Building cost versus damping cost Cost of damper as a ratio of total construction cost is approximately 0.15%. Even less if you consider that the tank is also required for fire suppression.

8.2.15.3.5 Building code The ASCE 7-10 (ASCE, 2010) and IBC (2012) were considered as the basic standards for the design of the building.

8.2 Mass Damping Systems Case Studies

8.2.15.3.6 Peer-reviewed project Walter P Moore was the peer reviewer for this project.

8.2.15.3.7 Design forces Seismic design: the seismic category A was considered. No particular seismic analyses were conducted since the seismic demand was much smaller than wind demand. Wind design: the maximum magnitude peak base moments due the extreme 700-year return ULS event were based on the wind tunnel study carried out by Windtech Consultants (Fig. 8.253). The estimated values are as follows: Mx: 2603 MNm, My: 2600 MNm, and Mz: 66 MNm. Refer to Fig. 8.254 for the structural axis convention.

8.2.15.3.8 Expected performance The estimated annual maximum combined peak acceleration, for the 1% inherent structural damping assumption and the assumed dynamic properties, was 9.1 mg. With the effect of the 2% additional damping (under normal SLS loading) from the liquid damper, the acceleration reduces to 5.9 mg, which provides a comfortable margin from the ISO 10137 (ISO, 2007) criterion of 9.3 mg. In

FIGURE 8.253 The Independent: wind tunnel test.

881

882

CHAPTER 8 Case studies of tall buildings

FIGURE 8.254 The Independent: structural axis convection.

addition, resultant 5-year return standard deviation acceleration reduces from 5.2 mg to 3.8 mg, compared to the ISO 6897 (ISO, 1984) criterion of 5.7 mg. Serviceability drift and deflection (10-year return) also reduced with the effect of the damper from 0.72 ft (0.220 m) down to 0.60 ft (0.183 m). SLS interstory drift is less than 1/2” (0.012 m).

8.2.15.4 Damping overview The damper is positioned at the center of the roof level of the tower and is designed to work in both orthogonal sway modes. A 9.1 m 3 9.1 m (in plan) tank is considered, with a height of 3.1 m (these are the internal dimensions of the tank). Furthermore, a notch with plan dimensions of 2.0 m 3 2.0 m is removed

8.2 Mass Damping Systems Case Studies

from the southeastern corner of the tank. The water depth within the tank is approximately 2.4 m, such that the total water volume is to be 189,271 L. The tank sits within the core walls on level 62. Energy dissipation elements are located within the tank, with specifications as follows: A pattern of crucifix and single vertical fins extending from the bottom of the tank to at least the top of the waterline. A submerged platform approximately 0.7 m below the top of the water level. Note that the final position of the submerged platform below the waterline will depend on the natural frequency of the building. Construction of the platform should allow it to be attached to the vertical fins at any level. Alternatively, the tank is to be fabricated only after Windtech has measured the natural frequencies on-site of the completed tower structure. A 15-cm gap needs to be provided between the edge of the platform and the inside face of the walls of the tank.

8.2.15.4.1 Damping strategy This unique design for a liquid damper is a kind of semituned damper (Figs. 8.255 and 8.256). When tuned, it can provide up to 2% additional damping to the structure but can also provide 0.9% damping at a tuning ratio of 0.8. The primary aim is to further improve occupant comfort under building motion. However, a conservative assumption of 0.5% auxiliary damping was adopted for the ULS case due to the ability of the damper to still provide some damping to the system even when it is not tuned to the natural frequency of the building. This is possible due to the unique design of this liquid damper. 4

5 Roof framing per plan

8' min T/STL per plan

Stiffening ribs per MFR

T/SLAB per plan Concrete slab and walls per plan typ.

CLR

6" Typ. (VARIES +/–8") Ref. No: E8

10"–1 1/2" 7'–9 1/2" 6'–2 1/2"

2'–0"

Water Horizontal plate connected to baffles plate elevation to be confirmed from site testing

Vertical baffles continuous through tank

1'–6"

FIGURE 8.255 The Independent: liquid damper section.

Section

Concrete curb per plan

883

884

CHAPTER 8 Case studies of tall buildings

FIGURE 8.256 The Independent: liquid damper plan.

8.2.15.4.2 Damping type The acceleration of water within a liquid damper is typically 2 times that of the building’s acceleration but can be as much as 3 times higher. With the inclusion of baffle screens (or other types of energy-dissipating devices) the effect is reduced, such that the acceleration of the water within the liquid damper is only twice as high as that of the building. The results of the wind tunnel study indicate that assuming the inclusion of an auxiliary damper system provides 0.5% additional damping, the peak building acceleration for the translational motions for the 700-year return period extreme wind is 63.5 mg (5.1 m/s2). Given that the total mass of water within the damper is 189,271 L. The lateral loads exerted by the liquid onto the tank walls is calculated using Newtown’s second law with the aforementioned building acceleration increased by a factor of 2 to account for the higher water accelerations. Hence the total peak force for the 700-year return period extreme wind event is 235 kN. This force can act in the X or Y axis directions.

8.2 Mass Damping Systems Case Studies

8.2.15.4.3 Structural and damping design Code and guidelines. IBC (2012), ASCE 7-10 (ASCE, 2010), ACI 318-11 (ACI, 2011), AISC 360-10 (AISC, 2010). For the wind tunnel and damper there were other codes listed in the wind tunnel report for the design of the damper tank. Analysis modeling and software. The damper performance was validated using a method similar to that of shake table which had an inherent system damping of 0.5%. A total of 10 tests were performed at two different scales, each for a range of natural frequencies. A sample output is shown in Fig. 8.257. Nonlinear analysis was performed to determine the contribution of the damper to the overall system. Design principles. The initial aim was to dampen building motion by the use of a TLD as an added benefit to bring accelerations well under the most stringent criteria with the effect of the additional 2% damping. An initial assumption of 2.5% of critical damping had been adopted for the 700-year event, in line with full-scale data of tall buildings, observed by Spence and Kareem (2013). However, following the peer review, the value for the inherent structural damping was revised down to 2.0% for the 700-year event and therefore required the use of the tank to gain the additional 0.5%. Hence the damper tank that was initially an added benefit to bring accelerations well under limits was relied on to provide an additional 0.5% damping. This involved moving away from a simple TSD design to the current design to ensure that the damper is not as sensitive to tuning and can achieve at least 25% efficiency during the extreme 700-year event. Design phase considerations. Initially considered a larger, stiffer core without damper tank. Used wind computer model (a.k.a. desktop study) to develop design through DD and set primary core sizes. Wind tunnel study confirmed the results of computer model and forces utilized for CDs.

FIGURE 8.257 The Independent: decay plot showing damper performance.

885

886

CHAPTER 8 Case studies of tall buildings

8.2.15.4.4 Architectural integration strategy The damper integrates well with the architecture given that this innovative design does not require any significant change to the overall tank dimensions, which is intended as a fire hydrant tank.

8.2.15.4.5 MEP integration strategy This passive liquid damper design has no power requirements.

8.2.15.4.6 Elevator and other mechanical devices integration strategy This damper is located above the top of the lift core. Access is via a set of stairs from the lower level to one corner of the room that houses the liquid damper, hence the notch in plan.

8.2.15.4.7 Experimental tests Extensive testing has been carried out for different frequencies and with various configurations with the aim of optimizing the design. This test program was also repeated for a different model scale and results were consistent between the two model scales.

8.2.15.4.8 Monitoring system No monitoring system is proposed as this design of damper is not as sensitive to tuning as a conventional sloshing tank. Also a conservative assumption was taken for the effect of the damper in the ULS design case.

8.2.15.4.9 Maintenance strategy The damper has been designed to allow access for maintenance and cleaning. This includes provision of externally flanged hatches as well as bolt down hatches on the horizontal platform. Also the tank will be constructed using SS 316 steelframed fiberglass panels, which requires minimal maintenance.

8.2.15.4.10 Development process The damper was initially to be designed as a series of sloshing tanks. However, the architect requested that the damper utilizes a single tank, which was not possible given the dimensional constraints. Windtech Consultants tested a number of ideas including a similar design to the final configuration but without the intermediate platform and all failed to generate any significant damping. Windtech Consultants then came up with the idea of the intermediate platform that would engage the lower compartment in the same way as a dashpot. This feature generated a significant boost to the performance of the damper.

8.2 Mass Damping Systems Case Studies

8.2.15.5 Comparison of different design strategies 8.2.15.5.1 Structural options considered The addition of the damper tank allowed DCI to reduce the thickness of the core walls by 152
203 mm. This change essentially provided an additional 465 m2 of additional salable residential area to the project.

8.2.15.5.2 Damping solution considered Only liquid damper solutions were considered given the advantage of being able to double up as a fire hydrant tank.

8.2.15.5.3 Cost
benefit analysis The building required the use of a fire suppression tank at the roof. It made sense to utilize the tank also for damping thereby saving concrete and rebar in the core, as well as giving 464.5 m2 of sellable area back to the owner on each floor plate. Damping cost. The cost of the damper is negligible especially considering that the tank will also be required for fire suppression. Design implications. The effect of the damper under SLS and ULS loads is summarized in Tables 8.42 and 8.43.

8.2.15.6 Lesson learned and recommendations The request to utilize a single tank as well as operate within tight dimensional constraints has forced Windtech Consultants to back to first principles and look at Table 8.42 The Independent: Performance Under Serviceability Load (With 1.0% Inherent Structural Damping)

Natural Frequency

Damping From Tank in Isolation ( %)

Additional System Damping Due to the Tank (%)

Total System Damping (Inherent 1 Auxiliary) (%)

0.15 Hz (full-scale) 0.15 Hz (full-scale) 120% 0.15 Hz (full-scale) 220%

10 . 10 4

2.1 . 2.1 0.9

3.1 . 3.1 1.9

Table 8.43 The Independent: Performance Under Ultimate Load (With 2.0% Inherent Structural Damping)

Natural Frequency

Damping From Tank in Isolation (%)

Additional System Damping Due to the Tank (%)

Total System Damping (Inherent 1 Auxiliary) (%)

0.15 Hz (full-scale) 0.15 Hz (full-scale) 120% 0.15 Hz (full-scale) 220%

10 . 10 4

2.0 . 2.0 0.9

4.0 . 4.0 2.9

887

888

CHAPTER 8 Case studies of tall buildings

the basic physics of how damping is generated to explore innovative ideas for a new type of liquid damper.

8.2.15.6.1 Difficulties in the design This was a challenging assignment given the dimensional constraints as well as the fact that the liquid damper needed to be in the form of a single tank. This has forced Windtech Consultants to investigate innovative ideas to achieve a substantial level of performance from a liquid damper given the design constraints.

8.2.15.6.2 Design innovative solutions This is the first time that a water tank has been designed to be able to provide significant damping over a substantial range of natural frequencies.

8.2.15.6.3 Possible improvements Although this design has gone through a number of iterations to improve efficiency, there is potential for significant further improvement through the optimization of certain parameters in the design.

8.3 BASE ISOLATION SYSTEMS CASE STUDIES In this section, a case study of a building equipped with base isolation damping systems is reviewed. As explained in the previous chapter this technology has not been applied extensively for tall buildings.

8.3.1 NUNOA CAPITAL BUILDING, SANTIAGO, CHILE 8.3.1.1 Project data The major building data (Fig. 8.258) are summarized as follows: • • • •

• • • • • • •

Year of completions: 2016 Developer/contractor: Empresas Armas (Chile) Architectural designer: Empresas Armas (Chile) Structural designer Rene Lagos Engineers (Structural Design) (Chile) Ruben Boroschek and Associates (Isolation System Design) (Chile) Damping supplier: Dynamic Isolation Systems Inc. (United States) Testing laboratory: Dynamic Isolation Systems Testing Lab (United States) 1 EUCENTRE (Italy) Height: 86 m Interstory height: 2.55 m Slenderness ratio: 1:2.7 Stories: 29 above grade 1 4 below grade Gross area of the tower: 40,000 m2

8.3 Base Isolation Systems Case Studies

FIGURE 8.258 Nunoa Capital Building: building overview.

• •

•

Building function: Residential Structural material Concrete: H40 (f0 c 5 35 MPa) for the first 15 stories, H35 (f0 c 5 30 MPa) for the next 6 stories, H30 (f0 c 5 25 MPa) for the last 10 stories; f0 c is the compressive strength according to ACI318-14 (ACI, 2014) Rebar: A630-420H with a yield strength of 420 MPa and ultimate strength of 630 MPa Structural system: Dual (RC core bearing walls 1 intermediate moment-resisting frames) Typical wall thickness varies from 60 to 40 cm Typical beam section is 50/60 cm (width/height)

889

890

CHAPTER 8 Case studies of tall buildings

•

•

Damping system: 24 devices, 16 of them have a lead core (LRB), while the 8 remaining do not have a lead core (RB). Type A: 8 LRB isolators 115 cm diameter with a 20,000 kN capacity Type B: 8 LRB isolators 135 cm diameter with a 30,000 kN capacity Type C: 8 NRB isolators 155 cm diameter with a 40,000 kN capacity Floor plan area: Each tower has approximately 500 m2 per floor

8.3.1.2 Introduction/history During the 2010 Mw8.8 Chile earthquake, and despite the general good structural behavior observed, a lot of nonstructural and content damages were observed. The earthquake caused more than US$30 billion in direct losses. The indirect losses remain unknown, but it is estimated that they far exceed the direct losses. In response to the extensive nonstructural damage in residential facilities and general business operation disruption, investors and stakeholders triggered an increased demand for the use of seismic protection technologies such as seismic base isolation and energy dissipation systems. This was the case of the Nunoa Capital Building, a new residential building designed shortly after the earthquake, where the owner wanted to evaluate the feasibility to incorporate different types of seismic protection technologies, including energy dissipation and seismic isolation systems, resulting the latest in the most cost-effective approach notwithstanding the building height.

8.3.1.3 Structural system Typical plan layout of the isolated towers is shown in Fig. 8.259 (in blue vertical elements, in yellow beams of perimeter moment frame). Wall thickness varies from 60 to 40 cm. Typical beam section is 50/60 cm (width/height). Instead, in Fig. 8.260, it is shown the typical plan layout for underground levels. Both isolated towers share undergrounds (in red). The surrounding underground structure (in blue) is not isolated. Between the isolated structure and its surroundings a 50-cm gap was provided (green) to prevent pounding between isolated and fixed base structures.

8.3.1.3.1 Building fundamental periods The fundamental period of the fixed base building was 2.30 seconds. With the isolation system, the effective vibration period for the design basis earthquake (DBE) became 5.34 seconds. A comparison between the modal shape of both the fixed base and the isolated configurations of the building can be seen in Fig. 8.261. It can be seen that, due to the high lateral stiffness of the building, resulting from the application of the Chilean seismic design standards, the isolation system can effectively concentrate on the displacement demand. For the other base-isolated modes the vibration periods and their respective participating mass ratios are given in Table 8.44. The seismic mass of the building considered 100% if the dead load plus a 25% of the live load resulting in a total seismic weight of 400,000 kN. The elastic

FIGURE 8.259 Nunoa Capital Building: typical plan layout of isolated towers.

FIGURE 8.260 Nunoa Capital Building: typical plan layout for underground levels.

FIGURE 8.261 Nunoa Capital Building: fundamental frequency with and without base isolation.

8.3 Base Isolation Systems Case Studies

Table 8.44 Nunoa Capital Building: Base Isolation Mode Shapes and Participating Mass Ratios Mode

Period (s)

Dir. X Mass Ratio (%)

Dir. Y Mass Ratio (%)

Rot. Mass Ratio (%)

1 2 3 4 5 6 7 8

5.34 5.26 4.83 2.13 1.55 1.41 1.40 1.13

98.25 0.56 0.04 0.00 0.03 1.10 0.01 0.00

0.56 98.61 0.00 0.00 0.53 0.01 0.00 0.28

0.04 0.00 98.65 0.05 0.00 0.01 0.68 0.00

FIGURE 8.262 Nunoa Capital Building: seismic isolation layout.

seismic base shear demand for the isolated building was equal to 13,500 kN, corresponding to 3.4% of the seismic weight. For isolated structures, the Chilean code allows for considering a response modification factor R 5 2 but imposes a minimum design base shear of 5% as long as it is not greater than the elastic demand. This resulted in a building designed directly with the elastic demand and therefore no structural damage for the DBE is expected.

8.3.1.3.2 Damping strategy utilized Twenty-four isolators (8 RB and 16 LRB) are used at the base of the building, distributed as shown in Fig. 8.262. They effectively decouple the isolated structure from the ground and provide both high flexibility and high damping.

893

894

CHAPTER 8 Case studies of tall buildings

8.3.1.3.3 Additional damping provided by the damping system The total system damping provided by the isolation system is 25% for the DBE and 20% for the MCE.

8.3.1.3.4 Building cost versus damping cost The isolation system cost is estimated to account for 0.75% of the cost of the condominiums.

8.3.1.3.5 Building code NCh2745:2013 (NCH, 2013) (Chilean Code for isolated structures based on American standard UBC-97).

8.3.1.3.6 Peer-reviewed project The project was peer reviewed by IEC.

8.3.1.3.7 Design forces Design of Chilean buildings placed in a high seismic zone is heavily controlled by seismic rather than wind forces. The elastic spectrum for the DBE is shown in Fig. 8.263. Vibration modes associated with the seismic isolation, with periods larger than 3 seconds, are reduced due to additional damping provided by the LRBs (25% for DBE).

FIGURE 8.263 Nunoa Capital Building: DBE elastic spectrum.

8.3 Base Isolation Systems Case Studies

8.3.1.3.8 Expected performance Interstory drifts for DBE are less than 0.3%. With such a low interstory drifts, a seismic performance of immediate occupancy is expected.

8.3.1.4 Damping overview A common seismic isolation system was implemented for the two towers, which are connected at the underground levels, and rest on top of a 2 m in thickness slab supported by 24 natural rubber isolators, 16 of which are LRBs. The isolators are mounted on footings connected to each other by beams. The isolated towers and the peripheral structures are separated by a 50-cm isolation gap which is significantly larger than the requirement of the Chilean isolation code NCh2745:2013 (NCH, 2013) in order to minimize the probability of impact between the isolated structure and the adjacent structures.

8.3.1.4.1 Damping strategy For the Nunoa Capital Building, the use of various energy dissipation devices was evaluated. These included viscous dampers, viscous wall dampers, VE walls, and TMDs in combination with viscous dampers. The use of viscous dampers was discarded during the first stages of the project, due to their architectural effect on the fac¸ades. For the same reason, the use of TMDs was also eliminated, as they would only be used in combination with viscous dampers distributed along building height. Consequently, in one of the first stages of the project, the use of viscous walls in nonstructural partitions was evaluated, as well as viscous walls coupled with the walls of the elevator shaft. These are shown schematically in Fig. 8.264. In addition, the feasibility of using a seismic isolation system in two configurations was evaluated: below each tower on the first floor and at the base of the entire structure below the lowest underground level. Fig. 8.265 shows a comparison of the seismic responses obtained for the four analyzed cases. This preliminary comparison was completed considering historical earthquake records that are compatible with the spectrum of the Chilean isolation code. The preliminary analyses completed indicated that the cost of implementing the energy dissipators in partitions and coupling walls was between US$2 and US$2.5 million, while the cost of the isolation systems required for isolating the towers individually or together was between US$700,000 and US$1.2 million. In light of the economical and technical analyses performed, it was quite evident that the most appropriate option for the structure was the use of seismic isolation systems.

8.3.1.4.2 Damping type Among the existing seismic isolation systems, the authors decided to use a combination of natural rubber and LRBs, because of the stability and predictability of their properties. Other isolation systems, such as frictional isolators, were discarded due to the difficulties they present for predicting and modeling the

895

FIGURE 8.264 Nunoa Capital Building: energy dissipation alternatives considered: (A) viscous walls in partitions (in red); and (B) viscous walls coupling concrete walls (in red boxes).

8.3 Base Isolation Systems Case Studies

FIGURE 8.265 Nunoa Capital Building: comparison of elastic seismic response of conventional structure and the structure with base isolation and energy dissipaters.

897

898

CHAPTER 8 Case studies of tall buildings

variation of the friction in the isolator during seismic movements, given the vertical effects of earthquakes. Similarly, the use of high damping rubber isolators was dismissed for the low level of damping they provide, and for the difficulties in predicting their behavior during severe strong motions. The seismic isolators of the building are located under the ends of the walls and under the columns (Fig. 8.267). The project uses 24 natural rubber isolators manufactured by Dynamic Isolation Systems. These devices are made of rubber with a strain capacity over 600%, whose long-term properties are extremely stable. The design process considered an extensive database of experimental results obtained for similar isolators, subjected to comparable vertical loads, and subjected to even larger displacement levels. Of the 24 devices, 16 of them have a lead core (LRB), while the 8 remaining do not have a lead core (RB). The largest diameter Type C (RB) isolators are 155 cm in diameter, have a load capacity of more than 40,000 kN, and they will be located in the most heavily loaded locations under the ends of the walls of the elevator shafts. There are eight type A LRB isolators, 115 cm in diameter, with a load capacity of more than 20,000 kN, and eight type B isolators, 135 cm in diameter, with a load capacity of more than 30,000 kN. The stiffest isolators are the type B isolators and they are located at the perimeter in order to control the torsion of the structure. The seismic isolation system achieves an effective vibration period near to 5 seconds and about 20% effective damping. Based on the results of the seismic analyses, reductions of shears, absolute accelerations, and interstory drifts between 70% and 80% are obtained compared to the fixed base elastic response.

8.3.1.4.3 Structural and damping design Code and guidelines. For the seismic design of the isolated structure and the seismic isolation system, the requirements of the Chilean standard NCh2745 (NCH, 2013) were fulfilled. The applicable requirements of NCh433 (NCH, 1996), for the seismic design of conventional structures, were also fulfilled. Due to the period of the fixed base structure, around 2 seconds, it was estimated that the period of the isolated structure would be around 5
6 seconds, and therefore, in accordance with the current Chilean isolation code, a site-specific seismic hazard study was required. Analysis modeling and software. In the Nunoa Capital Building, two types of analyses were used. First, a linear analysis, specifically a response-spectrum analysis, was performed. In this analysis, the isolators were modeled with their equivalent secant stiffness at the design displacement level. A site-specific design spectrum obtained from the seismic hazard evaluation was used. This spectrum represents the demand for the DBE level and was reduced by a factor of 2 for periods larger than 3.0 seconds in order to consider the damping provided by the seismic isolation system. This initial estimated factor of 2 was later verified through nonlinear analysis. This linear analysis was used to design the superstructure, the foundations, and to define the seismic isolation system. A nonlinear time history analysis was then carried out. In this analysis the isolators were modeled

8.3 Base Isolation Systems Case Studies

with a nonlinear biaxial hysteretic model that uses coupled plasticity properties for the two shear deformations. This causes the extra damping of the seismic isolation system to be directly considered and provides an estimation of the effective damping of the system. This analysis was used primarily to validate the preliminary design of the seismic isolation system. Seven pairs of ground acceleration records obtained from the seismic hazard evaluation were used. These ground acceleration records represent the demand for the MCE level. As allowed by the Chilean isolation code, all the parameters of interest such as displacements or axial loads on the isolators were obtained as the average of the peak response for each individual pair of records. As explained later, the superstructure is expected to remain essentially elastic for both the DBE and the MCE, which allows for modeling the building by using the linear elastic elements available in ETABS. Design principles. The principles of the Chilean code for seismic isolation consider that damage, either structural or nonstructural, are not acceptable for the DBE, defined as the ground motion with a probability of exceedance of 10% in 50 years. For the MCE, defined as the ground motion with a probability of exceedance of 10% in 100 years, failure of the isolation devices is not accepted although some extent of damage to both structural and nonstructural components is permitted. Design phase considerations. There were basically two design phases: a preliminary design where the feasibility studies of all the seismic protection strategies declared in Section 3.1 were made and a second design development phase which ended with construction design documents and open specification for the seismic isolators.

8.3.1.4.4 Architectural integration strategy The seismic protection alternative was selected in order to minimize its impact on the architecture of the building. For that reason, the use of seismic isolation under the underground levels resulted in a null impact on the architectural design of the towers. The only evidence of the existence of the isolation is the expansion joint cover placed all around the seismically isolated structure.

8.3.1.4.5 MEP integration strategy The greatest challenge on MEP integration on seismically isolated buildings is the large differential displacements that should be accommodated by each utility that cross the isolation plane.

8.3.1.4.6 Elevator and other mechanical devices integration strategy In order to prevent any impact of the protection system on the elevators system design, the isolation plane was chosen to be under the lowest level of the structure. The elevator’s pits fit perfectly in the 2 m in thickness concrete slab placed on top of the isolation system. Flexible hoses (Fig. 8.266) were considered for all utilities crossing either horizontally or vertically the isolation plane. Flexible

899

900

CHAPTER 8 Case studies of tall buildings

FIGURE 8.266 Nunoa Capital Building: example of flexible for fir extinguishing system.

elements were considered for drinking water, fire extinguishing system, residual and rain water, electrical and data ducts, and gas pipelines, among many others.

8.3.1.4.7 Experimental tests The seismic isolation system was subjected to an extensive test series carried out at Dynamic Isolation Systems’ laboratories in McCarran, Nevada, and Eucentre laboratories in Pavia. Prototypes and all production isolators were subjected to an exhaustive test sequence of combined compression loads and shear displacements in order to validate the properties assumed for the design and to verify the stability of the isolators under extreme seismic loads.

8.3.1.4.8 Monitoring system An extensive temporary instrumentation system comprising of several high sensitivity accelerometers was installed right after the building construction ended up. From environmental vibration measurements, the dynamic properties of the structure were determined, obtaining results similar to those predicted by the analysis models.

8.3.1.4.9 Maintenance strategy According to the requirements of the Chilean standard NCh2745 (NCH, 2013), a maintenance/inspection program was prepared by the authors. The program requests inspecting the seismic protection system whenever the structure is

8.3 Base Isolation Systems Case Studies

affected by a flooding, fire, or earthquake with modified Mercalli intensity larger than VI. The inspection shall be performed, as a maximum, every 5 years, in order to check the conditions of the isolation system and the nonstructural components and systems crossing the seismic isolation gap horizontally or vertically. The project did not consider fire protection for the isolators given that the isolators were installed in a mechanical floor with restricted access to people and without fire loads or hazards.

8.3.1.4.10 Development process The design team consisted on of two different specialists: Rene Lagos Engineers (RLE) for the structural design of the tower and Ruben Boroschek and Associates (RBA) for the design of the seismic protection system. Due to the expertise of the design team on seismic protection devices, no special interaction with different isolators manufacturers and laboratories was required until the final stage where the isolators supply was adjudicated.

8.3.1.5 Comparison of different design strategies 8.3.1.5.1 Structural options considered The isolation system greatly diminishes its performance for flexible buildings. In this specific case, in order to be technically feasible, the building needed to be as stiff as possible. This immediately discarded the possibility to use a structural system based on a moment frame. Instead, a strong RC wall core was considered surrounding the vertical circulations (stairs, elevators), which provided the building the stiffness it needed for the isolation system to effectively work.

8.3.1.5.2 Damping solution considered Additional damping solutions consisting of viscous walls were evaluated as explained in Section 8.3.1.4.1.

8.3.1.5.3 Cost
benefit analysis The isolation system cost is estimated to account for 0.75% of the cost of the condominiums. This cost includes manufacture, testing of prototypes, and 100% of production, and transportation to construction site.

8.3.1.6 Lesson learned and recommendations 8.3.1.6.1 Difficulties in the design The principal design challenges were the following: Obtaining low lateral stiffness while maintaining a high vertical stiffness of the isolation system (Chilean code (NCH, 2013) requires a minimum vertical frequency of 10 Hz). Controlling maximum compression forces and avoiding tension forces on the isolators for the MCE (Fig. 8.267). This was achieved in part by using a 2-m thickness slab (Fig. 8.268) just above the isolation system that stiffens and

901

902

CHAPTER 8 Case studies of tall buildings

FIGURE 8.267 Nunoa Capital Building: maximum and minimum force on isolator for MCE.

FIGURE 8.268 Nunoa Capital Building: 2-m thick slab above isolation system to couple external isolations with internal ones.

8.3 Base Isolation Systems Case Studies

29.00 26.00 23.00

Story

20.00 17.00

With outrigger

14.00

Without outrigger

11.00

Code limit

8.00 5.00 2.00 –1.00 –4.00

0

0.0005

0.001

0.0015

0.002

0.0025

0.003

0.0035

Interstory drift (%)

FIGURE 8.269 Nunoa Capital Building: maximum drift with and without outrigger.

FIGURE 8.270 Nunoa Capital Building: one-story height outrigger on the roof to control code strict maximum interstory drift limitation.

couples the isolation system so that the overturning moment is resisted by the whole plan rather than the isolators just below the RC core wall. Tension forces were completely avoided and maximum compression forces were roughly higher than 40,000 kN. Controlling interstory drifts. Chilean code has a very strong limitation for the interstory drifts (Fig. 8.269). Specifically for isolated buildings, the interstory drifts measured directly for the DBE forces shall be lower than 0.25%. This was achieved by using an outrigger in the roof coupling the core wall with the external columns (Fig. 8.270). Although it is not the optimal location for an outrigger, it effectively reduced the interstory drifts just below the code requirement without a large impact on the architecture.

903

904

CHAPTER 8 Case studies of tall buildings

8.3.1.6.2 Design innovative solutions This building is the tallest residential building in the America incorporating seismic isolation systems. This project demonstrates the technical feasibility of using seismic isolation in tall buildings in order to improve its seismic performance, as far as strict requirements on interstory drifts are imposed. The costs of the protection system are proven to be negligible in comparison to the project costs. This project also demonstrates that seismic protection technologies can be used not only for hospital or critical facilities but also for condominiums and offices, triggering an increasing demand for the use of seismic protection devices in buildings.

8.3.1.6.3 Possible improvements We strongly think that for this particular project and its special characteristics (structure stiffness, stiff soil, etc.), seismic isolation was the best alternative to achieve additional seismic protection.

8.4 ACTIVE, SEMIACTIVE, AND HYBRID SYSTEMS CASE STUDIES In this section, a case study of a building equipped with active system is reviewed. The application of this technological solution is not widespread worldwide but with the advancements of different active system in the near future it is expected to have more applications in tall buildings.

8.4.1 THYSSENKRUP TEST TOWER, ROTTWEIL, GERMANY 8.4.1.1 Project data The major building data (Fig. 8.271) are summarized as follows: • • • • • • • • • •

Year of completion: 2016 (Topped out) Developer/contractor: Krupp Hoesch Stahl GmbH/Thyssen krupp Elevators (Germany) Architectural designer: Werner Sobek with Helmut Jahn, Stuttgart and Chicago (Germany) Structural designer: Werner Sobek, Stuttgart (Germany) Damping supplier: GERB Vibration Control System (Germany) Testing laboratory: Wacker Ingenieure (Wind Engineering/Wind Tunnel Tests) (Germany) Height: 246 m Interstory height: 3.3 m Slenderness ratio: 1/11.8 Stories: 27 above grade/8 stories office space

8.4 Active, Semiactive, and Hybrid Systems Case Studies

FIGURE 8.271 Thyssenkrup Test Tower: building overview.

• • •

• • •

Gross area of the tower: 340 m2 Building function: Elevator test tower/offices/observation platform (232 m) Structural material Reinforced concrete (C50/60) Polytetrafluoroethylene (PTFE)-membrane fac¸ade on a steel structure Damping type: Hybrid mass damper system (240 ton) Floor plan area: 20.8-m diameter circular plan Structural system: Cantilevered reinforced concrete core on a raft foundation (32 m)

8.4.1.2 Introduction/history A 246-m tall tower has been erected in South West Germany with the purpose to accommodate a testing facility for elevator innovations. The tower has a circular footprint with a diameter of 20 m and provides space for nine testing elevator shafts, a fire service elevator and a glass panorama view elevator. In addition a

905

906

CHAPTER 8 Case studies of tall buildings

220-m shift is provided to be used as mechanical lifting shaft to ensure the test programs. The tower base has a diameter of 40 m to provide additional space for service facilities, a lobby and an education center. At a level of 232 m an observation platform—Germany’s highest—allows a spectacular view to the Black Forest and even the Alps on clear days.

8.4.1.3 Structural system The structural system of the Test Tower is basically an RC tube with a diameter of 20.8 m (Fig. 8.272) that is clamped 30 m into the surrounding soil. Up to a level of 110 m the tube has a thickness of 40 cm and above of 25 cm (Fig. 8.273). The soil, consisting of a Keuper layer and subjacent shell lime stone, provides a high bearing capacity so a pile foundation was not required. In addition to the clamping by the surrounding soil, the wider tower base structure provides additional lateral stiffness. The internal stiffening of the concrete tube results mainly from the inner walls of the elevator shafts. Only at certain height levels prefabricated ceiling slabs were installed to allow an access to the lift shafts. Some of the lift shafts end at a level of 115 m and are covered by 40 cm thick long span floors which were casted in place. These floors are used for office space. Above a void space up to

FIGURE 8.272 Thyssenkrup Test Tower: plan view.

8.4 Active, Semiactive, and Hybrid Systems Case Studies

FIGURE 8.273 Thyssenkrup Test Tower: structural elevation.

a level of 197 m is used as a heat reservoir and for the TMD system. The top of the tower is used for office space and for the remaining lift shafts. A distinctive feature of the test tower is the textile fac¸ade which consists of a PTFE-coated glass-fiber mesh. The aperture width of the mesh increases with the height of the building to increase translucency of the fac¸ade, decrease the density and weight of the material and for aerodynamic effects. The spiral shape of the fac¸ade is meant to function as a Scruton spiral and the fabric itself helps to shade the concrete structure to avoid thermal-induced stresses due to solar radiation. For the fac¸ade design and the choice of material aspects of the installation, the maintenance and durability as well as the wind loading had to be considered.

8.4.1.3.1 Building fundamental periods The fundamental frequencies for the Test Tower were predicted to be in range of 0.17
0.20 Hz seconds depending on the construction stage and state of the concrete (cracked/uncracked), see Fig. 8.274.

907

908

CHAPTER 8 Case studies of tall buildings

FIGURE 8.274 Thyssenkrup Test Tower: fundamental frequencies and mode shapes.

8.4.1.3.2 Damping strategy utilized To reduce the dynamic response to crosswind excitation, a passive TMD system was implemented. Since the tower shall be used as a test tower for building sway sensitive elevator equipment, the owner was looking for a possibility to artificially excite the tower on calm days, though with displacements that would not cause any fatigue issues. The requested level for the artificial sway motion was in the considered safe range of approximately 6 200 mm. This request gave the rare opportunity to implement a hybrid mass damper (HMD) or, since the design system deviates from other HMD systems that were introduced in the past, a socalled dual-use TMD. The system shall be presented in the following chapters, including the optimization of the passive system, the design of the actuators’ mechanism, control algorithms, and the safety concept.

8.4.1.3.3 Additional damping provided by the damping system The TMD works in a passive and an active mode. For the mode the supplementary damping that achieved by the TMD operation is about 3%
3.2%. For the active suppression mode of the HMD system the structural damping can be increased even further to the theoretical value of 4.5%.

8.4 Active, Semiactive, and Hybrid Systems Case Studies

8.4.1.3.4 Building cost versus damping cost Conventional measures to achieve the serviceability requirements would have resulted in a very significant increase of material and self-weight of the tower so the application of a TMD system was not only far more cost-effective but also least invasive for the structural system. Furthermore the owner of the tower wanted to implement an active system as an additional test feature.

8.4.1.3.5 Building code The tower was designed according to European Standards (EC 1 (CEN, 2010), EC 2, EC 3).

8.4.1.3.6 Peer-reviewed project According to the German standards the design was peer reviewed by a inspecting structural engineer (Breinlinger Ingenieure).

8.4.1.4 Damping overview 8.4.1.4.1 Damping strategy The duality of the control system serves (1) the purpose of an reduced energy consumption, while in normal passive operation, to reduce the occurring vibrations; (2) the purpose of reduced force requirements for the actuators, by using resonance effects in order to excite the main TMD mass for ultimately achieving the actual force demand. So, compared to other control systems (see Section 8.4.1.5.2) the actuators connect the main structure and the TMD mass but are not used to control the TMD mass directly as it would be the case for a typical active mass driver/damper system.

8.4.1.4.2 Damping type For the implemented dual-use TMD a reaction mass for the passive operation of 240 tons was chosen. For the excitation operation mode two linear drives—one in each principal direction—are attached to the TMD mass with pivots near the center of gravity of the mass to avoid any torsional artifacts. Each linear drive can provide forces up to 40 kN within a stroke of maximum 6 600 mm. The linear drives can be detached so that the entire passive mode will not be influenced by the bearings of the actuators for the unlikely event of a bearing failure (Fig. 8.275).

8.4.1.4.3 Structural and damping design The wind analysis (Fig. 8.276) revealed that resonant excitation of the structure can occur at wind speeds which correspond to ground values (height 10 m) in the range of 55
60 km/h. It was expected that without additional damping this resonant excitation would cause top deflections of about 6 750 mm which would not

909

910

CHAPTER 8 Case studies of tall buildings

FIGURE 8.275 Thyssenkrup Test Tower: (A) pendulum rope supported TMD mass and (B) linear motor as actuator.

FIGURE 8.276 Thyssenkrup Test Tower: structural model.

8.4 Active, Semiactive, and Hybrid Systems Case Studies

FIGURE 8.277 Thyssenkrup Test Tower: wind tunnel test model.

only cause discomfort for the occupants but would also have significant fatigue life implications for the concrete structure of the tower (Fig. 8.277). Code and guidelines. To assess the serviceability and to define the requirements for supplementary damping the ISO 10137 (ISO, 2007)—bases for design of structures—serviceability of buildings and walkways against vibrations was applied. Analysis modeling and software. The TMD has been discretely modeled as a pendulum system capturing also its eccentric position at the tower. The load characteristics for the governing input case, i.e., resonant excitation due to vortex shedding, are very alike to that of a single harmonic excitation. Yet, the coexisting gust loading is inherently of stochastic nature and for it, optimization criteria other than the well-known Den Hartog (1956) criterion apply. In addition, the relative displacements of the TMD mass are bigger for a stochastic than for a harmonic-type loading. Since for the numerical determination of the optimum TMD parameters a close to reality loading should be considered, a time history was generated that included both the stochastic gust loading (based on the Davenport Spectrum) and a superimposed resonant, vortex shedding like, component for representing the overall crosswind excitation (Fig. 8.278). The tower

911

CHAPTER 8 Case studies of tall buildings

100

2 1.5

10–1

1 0.5

10–2

0 10–3

–0.5 –1

10–4

–1.5 10–5

–2 0

500

1000

1500 Time (s)

2000

2500

1

0

0.2

0.4 0.6 Frequency (Hz)

0.8

1

0.8 Deflection TMD mass

with TMD w/o TMD

0.8

0.6

0.6

0.4

0.4

0.2

0.2

(m)

Tower deflection (m)

912

0

0 –0.2

–0.2 –0.4

–0.4

–0.6

–0.6

–0.8

–0.8 0

100

200

300

400 500 Time (s)

600

700

800

0

100

200

300

400 500 Time (s)

600

700

800

FIGURE 8.278 Thyssenkrup Test Tower: generated time history for the crosswind loading and resulting FFT spectrum (above), and tower displacement with and without optimized TMD and resulting TMD displacement (below).

deflection reduction that can be achieved with the optimized passive TMD system as well as the resulting TMD displacements is shown in Fig. 8.278. Based on these results for an estimated inherent structural damping of 0.8%, it was determined that a TMD mass of 240 tons was required to keep the displacements within 6 650 mm while maintaining an optimum TMD damping ratio for the best TMD performance. An increase in the TMD damping could have reduced the travel while the efficacy would still have been sufficient but this would have adversely affected the actuator force requirements. To determine the required forces for the optimum 240-ton TMD setup the analogous model has also been used to verify that, with a maximum force of 40 kN from the actuators, tower deflections can be achieved in the range of 6 200 mm. Design principles. The parameters of the passive TMD system had to be determined considering three different aspects: (1) to provide sufficient additional structural damping in order to reduce the dynamic response owing to vortex shedding excitation, (2) to limit the resulting TMD main mass travel in the passive mode, to an attainable/practical value, for when under gust crosswind excitation, and (3) to choose the TMD mass according to the energy input that is required

8.4 Active, Semiactive, and Hybrid Systems Case Studies

for the desired maximum tower deflection in the excitation mode, considering the performance envelope enabled by the provided actuators (i.e., maximum force generated and maximum stroke during operation). To optimize the TMD system a numerical model was used that represented the mass distribution of the tower, and mass moments of inertia as reported in the identified structural properties. The stiffness elements between the floors were also tailored to match the mode shapes and natural frequencies from full-scale observations. Fig. 8.279A shows the mode shapes and natural frequencies of this employed analogous model. Fig. 8.279B also compares the mode shapes of the analogous model against these of the detailed model prepared by the structural consultant. Design phase considerations. The main design goal was to cautiously integrate the tower into the cultural landscape around the medieval town of Rottweil. The key element for this is the helical membrane fac¸ade which give the tower a mild appearance. From a technical aspect the tower design had to ensure the accessibility to the elevator test shafts. Also the observation deck that was desired by the general public in Rottweil had to be esthetically integrated.

(A)

(B) 245 m

Mode 1

Detailed model Analogous model

0m 0

0.2

0.4

0.6

0.8

1

1.2

Normalized displacement 245 m Mode 2

Detailed model Analogous model

0m 0

0.2

0.4

0.6

0.8

1

1.2

Normalized displacement

FIGURE 8.279 Thyssenkrup Test Tower: relevant modes and calculated natural frequencies for (A) model calibration and (B) mode shape comparison.

913

914

CHAPTER 8 Case studies of tall buildings

8.4.1.4.4 Architectural integration strategy Architecture and structural design complement one another in the design of the tower. The helical membrane fac¸ade reduces the wind and temperature-induced loads, so besides the esthetic aspects it also serves important technical aspects.

8.4.1.4.5 MEP integration strategy Noteworthy is the implementation of a heat reservoir. Since some of the elevator tests shafts end at a height level of 110 m while the cylindrical shape of the tower was retained all the way to the top, a vast space that is not occupied can be used for that. All MEP facilities are located in that space and the heat discharged from these facilities as well as from the elevator operation is stored in the reservoir to be reused when required.

8.4.1.4.6 Elevator and other mechanical devices integration strategy The tower serves as an elevator test tower. So besides an firefighter elevator and a scenic panorama to access the observation platform, 10 additional elevator shafts are provided to install and test elevator systems.

8.4.1.4.7 Experimental tests Initial vibration tests of the tower were performed at the current stage where the tower is not completed yet so the fundamental frequencies of the tower at this stage are above the specified tuning range of the passive TMD system. The TMD is adjusted to the highest possible tuning frequency. Objective of the vibration tests was the determination of the fundamental frequencies of the tower with blocked TMD system as well as the inherent structural damping. In addition the dynamic behavior with engaged passive TMD system should be determined as well as the increase in structural damping due to the passive TMD system. To identify the fundamental natural frequencies of the tower the averaged normalized power spectral density method (Wenzel et al., 1991) can be used. Fig. 8.280A shows the recorded time histories of the horizontal ambient vibrations in X and Y directions with locked TMD. Fig. 8.280B shows the resulting averaged auto power spectra for a segment length of 120 seconds. The spectra show that the tower shows a dynamic response at two dominant frequencies (0.225 Hz in X and 0.245 Hz in Y direction). Further to the above-described averaged power spectrum method—which assumes that the ambient excitation causes a sufficient dynamic response in the vibration modes of interest to gain stochastic security—the natural frequencies were determined using the commercial signal processing software ARTEMIS (Wenzel et al., 1991) which incorporates enhanced frequency domain decomposition and stochastic subspace identification methods. In addition to the ambient vibration tests with the passive TMD system, preliminary tests with the active excitation mode were performed despite the detuned state of the TMD system. Fig. 8.281A shows the time history of the recorded

8.4 Active, Semiactive, and Hybrid Systems Case Studies

(A) 5

(B)

× 10–3

1.4 NS direction OW direction

× 10–7 NS direction OW direction

1.2

(m/sx)x)

(m/sx)

1 0

0.8 0.6 0.4 0.2

–5

0

2000

4000

6000

8000 10000 12000 14000

Time (S)

0

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8 0.9

1

Frequency (Hz)

FIGURE 8.280 Thyssenkrup Test Tower: time history for the measure acceleration at the top of the tower in the two main directions (A) and corresponding APS spectra (B).

FIGURE 8.281 Thyssenkrup Test Tower: time histories of the occurring accelerations at the tower top during the artificial excitation (marked green: sinusoidal excitation/marked blue: controlled excitation mode with suppression in one of the main axis (A) and corresponding FFT spectrum (B).

accelerations at the tower top. After an initial sinusoidal excitation the algorithm for a steady acceleration level was enabled to experimentally determine the correlation between tower deflection and forced displacement of the TMD mass. The time history shown in Fig. 8.281B displays the two test scenarios in the excitation mode and the corresponding FFT spectra shows the clear response of the tower in its two fundamental frequencies (X and Y directions). The green marked time section is a sinusoidal excitation of the mass with disabled active control which caused a dynamic response in both main directions. The resulting FFT spectrum displays the two peaks to identify the fundamental frequencies in these two directions. The blue marked time segment is for an excitation with enabled active control which caused a more steady tower displacement dominantly in one direction. The resulting vibration decay after the excitation mode was switched off was also

915

916

CHAPTER 8 Case studies of tall buildings

used to determine the structural damping and determined to be 2.4%. This result correlates with the damping ratio determined with the soil
structure interaction algorithms.

8.4.1.4.8 Monitoring system The system is instrumented with four uniaxial accelerometers (seismic K-Beam/ MEMS; one in each direction) to determine the tower top level and the TMD mass accelerations. The acceleration signals are getting bandpass filtered inside the frequency range of the towers fundamental natural frequencies (0.1
0.3 Hz) and integrated to give the tower velocities and deflections. In addition, the integrated displacement values can be compared against the ones of a complimentary Global Navigation Satellite System installed also at the top to compare for signal drifts. An initial correlation test was performed accordingly. Furthermore, the TMD displacements are monitored directly with string pot transducers and an inductive length measuring system integrated within the linear motors.

8.4.1.4.9 Control strategies The general control loop for the tower is shown in Fig. 8.282 which illustrates the closed loop control sequence. A detailed design of the control algorithm would require analytical models of the actuators
TMD mass interaction to determine Closed loop

Disturbances

Nominal value + O–

Controller

Correction value

Actual value Actuator system

Tower displacement

Sensor tower

Controller

Sensor tower

Sensor TMD

Controller

Tower displacement

TMD displacement

Force output actuator

FIGURE 8.282 Thyssenkrup Test Tower: feedback control loop practice.

8.4 Active, Semiactive, and Hybrid Systems Case Studies

the connection between actuator input to applied action force. These should combine with the analytical model of the main structure that describes the overall building dynamic behavior. Further, compared to standard control algorithms that were applied to other HMD applications (Section 4.3, Chapter 4), the control algorithm for this tower application had to consider synchronous controlled excitation in one main direction and vibration reduction in the perpendicular direction. Since the dynamic response is expected to be relevant in the fundamental modes, linear feedback control with all its benefits to avoid instabilities could be applied for both these tasks. For the simple control realization practiced, the linear feedback control, which attenuates the towers’ dynamic response perpendicular to the excitation, drives the relevant actuator through a specified weighted linear sum of seven structural dynamic measurements. These measurements are TMD and tower top accelerations, TMD and tower top velocities, TMD and tower top displacements, and TMD to tower top relative displacement. It is important to note that the preselected weighting factors are simple, positive, or negative scalar gains. The parameterization of the weighting factors is based on the dynamic tests of the tower and are not using any frequency-dependent modification. The active actuator feedback is calculated instantaneously (at 50 Hz), and it is again a linear combination of the previous dynamic measurements. There is no nonlinear manipulation being used. The control design was also implemented in a time domain numerical simulation (i.e., direct integration) to evidence the performance of the actuator control vibration mitigation. For the excitation mode the same control approach has been used in combination with a displacement offset that represents the excitation of the tower to the desired displacement value. The offset is a sinusoidal function based on the detected fundamental frequency of each direction. The control output to counteract the variation of the top displacement due to other disturbances is then modulated on the sinusoidal offset function and the control value is adapted accordingly. The linear feedback control for reducing the towers dynamic response perpendicular to the excitation allows the force command to the actuator to be a specified weighted linear sum of the seven dynamic measurements of the program. The seven measurements are TMD and tower acceleration, TMD and tower velocity, TMD and tower displacement, and TMD to tower relative displacement.

8.4.1.5 Comparison of different design strategies 8.4.1.5.1 Structural options considered To allow a short construction time, separate casting methods were avoided. Instead a sliding form work was used, which allowed to erect the towers’ shaft in less than 4 months.

8.4.1.5.2 Damping solution considered Since the tower is basically a concrete shaft, discrete damping options (outrigger damping/wall dampers) are not applicable, so the choice of damping solutions is

917

918

CHAPTER 8 Case studies of tall buildings

rather limited. Also the heat reservoir offers a vast space to implement a TMD system. To save costs prefabricated concrete slabs were used for the effective mass.

8.4.1.5.3 Cost
benefit analysis As mentioned the implementation of a passive TMD system was the least invasive and therefore most cost-effective way to maintain the serviceability of the building. The owner invested in addition to an active system which also allows a monitoring of the building vibrations so changes of the structural system over time can be monitored. This information is an investment for the future since it enables to perform lifetime assessments with much higher certainties. Damping cost. The costs for the passive damping system are in the range of 1.5% of the total costs for the building. The active control mechanism together with the control units and the building monitoring system including the GNNS Module added up to a similar amount.

8.4.1.6 Lesson learned and recommendations A dual-use TMD has been installed at the 246-m tall Thyssen krupp Test Tower with the objective to purposely excite the tower to a controlled dynamic response in its fundamental frequencies. The objective of this excitation is to achieve a defined building sway in the two main directions of the tower. Based on numerical calculations it was found that a 240-ton TMD mass was required to achieve the required supplementary damping for the passive mode and to generate the required control force to achieve a tower top displacement of 6 200 mm. To create a steady displacement level for the excitation mode and to suppress the displacements caused by wind and the forced vibration component in the perpendicular direction, a control algorithm was developed and tested with numerical simulations. After the installation of the passive pendulum-type TMD system for which prefabricated concrete slabs were used as reaction mass, the actuators were commissioned and initial tests have been performed. The tests revealed that the inherent damping of the tower was higher than anticipated but all relevant modes could be determined clearly. The active excitation mode of the building is working and could be tested although not to the full extent to derive all operation parameters since the building is not entirely completed yet. Due to the premature state of the building the passive TMD could not be adapted to the determined fundamental frequencies. As soon as the TMD system has been adapted, further tests of the building will be performed that will also include tests regarding amplitude-dependent damping and the integrity of the safety concept.

8.4.1.6.1 Difficulties in the design One of the main difficulties for the design phase was the consideration of the several construction stages—for each stage a design analysis had to be performed. The most critical stage was when the tower was completed but the pedestal building was not yet erected, so the additional stiffness was not in place yet. As a

8.4 Active, Semiactive, and Hybrid Systems Case Studies

consequence, a complex analysis of the soil behavior with this additional load had to be performed.

8.4.1.6.2 Design innovative solutions The distinctive element of the tower is the helical fac¸ade of the tower. The challenge for the fac¸ade design was not only the installation planning but also to design according to wind loads and to maintain a weather resistance. A textile fac¸ade with dimensions like the one applied at the Test Tower in Rottweil is a novelty and represents an innovation in terms of lightweight design elements (PTFE-coated glass-fiber fabric) and technical features. One of the important features is the permeability of the fac¸ade fabric which increases with the building height. So it allows transparency for the office floors and reduces the wind loads. Wind tunnel investigations have been performed to investigate the positive effect of the fac¸ade permeability to reduce vortex shedding-induced vibrations.

8.4.1.6.3 Possible improvements The HMD allows further research on control algorithms to reduce wind-induced vibrations for tall buildings.

919

CHAPTER

Future of dynamic modification systems

9

CHAPTER OUTLINE 9.1 Improved Modeling of Structural Behavior .........................................................922 9.1.1 Improved Understanding of Actions on Structures .............................922 9.1.2 Improved Behavior Model of Materials and Devices ...........................923 9.1.3 Enhanced Computing Power ............................................................923 9.1.4 Improved Design Codes...................................................................924 9.2 Implementing Results of Technological Progress ...............................................925 9.2.1 Development of Existing Devices and New Hybrid Systems ................925 9.2.2 New Technologies ...........................................................................926

In previous chapters, the current state of the art of dynamic modification systems and their design for tall building applications have been deeply discussed. The developments in this field have been mainly achieved in the last 50 years due to the extensive research projects that have been carried out on the subject. As emerging dynamic modification technology continues to evolve, the future of dynamic modification devices in tall buildings is destined to provide further improvement in terms of dynamic performance of buildings. With time, the use of dynamic modification devices will only increase as society becomes more aware of the benefits of these systems in structures. In other areas of human design, intentional use of materials that dissipate energy has long been embraced. Present day automobiles cannot be imagined without shock absorbers to improve ride comfort. A vast array of foams make furniture such as sofas and beds more comfortable and less “bouncy” than in past generations. Where the damping means are still entirely distinct from the materials used for structural supporting purposes, the size and cost is nowadays under constant downward pressure. As members of the profession (owners, architects, engineers, etc.) become more accepting and aware of resilient design and performance-based design, high performance damping systems will become important tools for the world’s best engineers as they are building components that can be used to improve the performance of the structure. In addition, building codes across the world are starting to incorporate guidelines for the design of building structures with supplemental dampers and isolation devices (as discussed in Chapter 5). As the general structural engineering community has more Damping Technologies for Tall Buildings. DOI: https://doi.org/10.1016/B978-0-12-815963-7.00009-9 Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved.

921

922

CHAPTER 9 Future of dynamic modification systems

reference points and guidelines to increase the understanding of the requirements of designing with these types of system, the adoption of these systems will increase further. For example, for wind design there is still apprehension in utilizing dampers to reduce wind loads for strength design, even though dampers are typically the most rigorously tested and engineered components in the entire structure. It is expected that this will change, as more high-performance dampers become more widely accepted and understood. For these reasons, the scope of this chapter will be to discuss the possible future trends in dynamic modification systems for tall buildings. Looking at the source of progress, these developments can be grouped as: 1. Improved modeling of structural behavior 2. Implementing results of technological progress The authors hope that this book will be continuously updated, to reflect the new advancements in the field.

9.1 IMPROVED MODELING OF STRUCTURAL BEHAVIOR The aim of any design activity is to define a safety envelope for the future product—construction in this case. To do this, the designer applies the actions on the model of the structure and receives a response. The more accurate these models are, the closer the envelope is to reality and the more economical the proposed solution is. From this point of view, looking to future developments, the following expected improvements in the field of dynamic modification devices can be noticed: • • • •

Improved understanding of actions on structures Improved behavior model of materials and devices Enhanced computing power (computers and programs) Improved design codes

9.1.1 IMPROVED UNDERSTANDING OF ACTIONS ON STRUCTURES Improvement of the understanding of the building behavior with dynamic modification systems strongly depends on the available recording data for each type of load. These are utilized to determine the device nominal values and have an important influence on a better calibration of the partial safety factors. An important aspect for the development of dynamic modification technology is the need of a database (Tamura, 2009) that collects the behavior and seismic/wind response of buildings with these devices. For example, for wind response, several aerodynamic databases have been developed, such as: NIST (National Institute of Standards and Technology in collaboration with the

9.1 Improved Modeling of Structural Behavior

University of West Ontario), NatHaz (University of Notre Dame) (Zhou et al., 2003; Kwon et al., 2008) e-wind (Tamkang University) (Cheng et al., 2008), and TPU (Tokyo Polytechnic University) (Quan et al., 2007). These databases would be beneficial to understand the effectiveness of wind response for buildings equipped with dynamic modification devices. Similar approach could be carried out through building monitoring during earthquake events (Isyumov et al., 2010) (e.g., California Strong Motion Instrumentation Program (CGS) (Shakal et al., 1988)). All these available databases allow a better management of the randomness of seismic and wind actions and eventually a more accurate model of the building behavior. In addition, building health monitoring could become an essential tool for understanding the effectiveness of dynamic modification system in building response. As the utilization of monitoring devices will increase, a wider database of building response can be developed. Moreover, it is important to allow the worldwide access and sharing of all these resources, since they will be valuable for the enhancement of the whole field.

9.1.2 IMPROVED BEHAVIOR MODEL OF MATERIALS AND DEVICES Dynamic modification devices are no longer a result of the construction process; they are the product of an industrial activity, benefitting from all the quality control procedures available in industry. As a result, the dispersion of the results decreases and the behavior of materials and devices can be better reproduced by models used for structural analysis. One of the possible future outcomes in improving dynamic modification devices behavior is the possibility of control of multimodal building response. Higher modes can become predominant in the design of high-rise buildings and the utilization of dynamic modification devices could be used to reduce the influence of these modes (Tsai and Chang, 2002). One possible solution can be the utilization of multi-TMD (Daniel and Lavan, 2015) or -TLD (Saha and Debbarma, 2017) control systems. For these reasons, it is important to study the whole building behavior in order to understand the influence of higher mode effects.

9.1.3 ENHANCED COMPUTING POWER Modern-day modeling software coupled with the availability of nonlinear dynamic modification properties has enabled building designers to optimize building performance without sacrificing structural integrity, by using energy-absorbing elements. Several widely used software packages have enabled the designer to incorporate dynamic modification devices relatively easily. These finite element analysis software packages have already incorporated various damping mathematical models to allow the link elements to behave in either a linear or a nonlinear fashion by providing the possibility to define the major variables (e.g., damping

923

924

CHAPTER 9 Future of dynamic modification systems

coefficient and the desired exponent for viscous dampers). Another type of damping model available in some software packages is a bilinear one, which allows the damping function to abide by one damping coefficient until a given force threshold is met and then the damper will follow the linear curve of the second function using the second damping coefficient. These types of programs can also allow the user to specify the amount of damping along each local axis or to implement a general amount of stiffness or damping if the link element orientates beyond the specified local axis for the given damping properties. These software packages can even go as far as assigning a rotational inertia to individual local axes of the link element to account for the rotational resistance of the damper in mechanisms or stiffer connections. All these individual parameters within the damping models allow the engineer to accurately represent the dynamic characteristics of the system within the overall structural model. As these software packages will keep advancing in their mathematical modeling capabilities and increasing their computational power, there will be a direct translation of the design to hone an optimized solution with high precision while also giving the opportunity to design the structures’ failure sequence in a systematic manner. Additionally, building information modeling allows engineers and architects to share the same model, evolving with time during the concept and design stages, in order to drastically reduce cost and improve accuracy. This will also accommodate even more futuristic building designs with performance-based attributes. Performance-based design with optimal damping will become commonplace, thereby resulting in buildings that are safer, more comfortable, and less expensive to design and maintain. At the same time, dynamic analysis capabilities have been developing; they started off as linear modal methods and then progressed by allowing only some elements to behave in a nonlinear manner. This was subsequently recognized by the building codes as computers increased their computational powers. Eventually, this led to the advancements of dynamic time history analyses of structures, while also being able to handle full nonlinear capabilities of plastic deformations throughout the structural elements. The accelerated development of calculation devices (increase speed of calculation and of volume of data that can be analyzed) allows the development of more accurate computer programs, able to incorporate dedicated modules that reproduce the behavior of such devices. Dynamic modification device producers one day could deliver subroutines for modeling the device behavior in structural analysis programs, making the analysis even more reliable.

9.1.4 IMPROVED DESIGN CODES The introduction of new dynamic modification categories in codes and national standards and the development of their design procedures can now be noticed all over the world. As reviewed, in Section 5.1, just few devices (e.g., viscous, hysteretic, base isolation) have been prescribed in codes until now. However, in the near

9.2 Implementing Results of Technological Progress

future, it would be essential to have code-based prescribed criteria for other device categories in order to allow a wider utilization of these solutions in buildings. Some building codes now outline a performance criteria objective that a structure is required to achieve for a given amount of incurred damage so that the state of the structure can be categorized according to risk levels as “immediate occupancy,” “life safety,” or “collapse prevention,” for example. The advent of these codes first accepting the use of performance-based design from practicing engineers using dynamic time history analysis has allowed the use of dynamic modification devices to be implemented into models. Recently, for example, the Canadian National Building Code (NBC, 2015) has recognized the use of supplemental energy dissipating devices, instead of relying on adding more structural stiffness throughout the building by increasing the members’ sizes, to be able to counteract the story lateral forces. Allowing devices to dissipate energy as their main function, instead of requiring a structural member to dissipate energy as a secondary function after fulfilling its main function as a load-carrying member, demands a new approach compared to the traditional one. This way of thinking has been well established in other industries such as the automotive sector, which has provided suspension system with shock and vibration absorbing elements for decades.

9.2 IMPLEMENTING RESULTS OF TECHNOLOGICAL PROGRESS The present day technological exponential development has an important impact on the dynamic modification device industry too and the following expected improvements can be noticed: • •

Development of existing devices and new hybrid systems New technologies

9.2.1 DEVELOPMENT OF EXISTING DEVICES AND NEW HYBRID SYSTEMS Combination of different dynamic modification technologies and devices with variable properties could become an improved design solution to reduced uncertainty and resist different loading scenarios. Some of the future possibilities could be seen from these examples: •

•

Hybrid combination of base isolation and building connection (Kasagi et al., 2016). This would provide the resistance for different sources of uncertainties, therefore, increasing the building resilience (Takewaki, 2006, 2013; Takewaki et al., 2012). Utilization of multi-TMD (Daniel and Lavan, 2015) for improving damping on several vibration modes.

925

926

CHAPTER 9 Future of dynamic modification systems

•

•

•

•

Integration between combined wind and seismic control of damping devices. One possible solution in this direction would be a more widespread utilization of smart damping approaches: active and semiactive damper systems (Spencer, 2002). Active and semiactive systems advancements could consider control algorithms to optimize energy consumption and the development of alternative maintenance-free actuator. Moreover, with the introduction of housing electric energy storage, the problem with the reliability of the grip supplies energy (e.g., possibility of blackout during an extreme event) could be overcome. Adaptation of linear, nonlinear, and multistaged behavior to cope with different load scenarios. While viscous damping devices usually display a nonlinear behavior which has disadvantages for normal operation stages, it can be beneficial for extreme loads. Alternatively, viscoelastic elements provide a nearly linear behavior but, for the extreme load cases, higher damping coefficients would be beneficial. A lot of research and product development is done to optimize the behavior for special applications and to allow an increased performance. Heat dissipation, larger strokes, etc., require additional devices so the conventional dampers can achieve. Fluid viscous dampers’ near-future progress can be envisaged, as the development of completely passive devices able to change their force versus velocity relationship (see Chapter 4) in different ranges of velocity, that is, able to change the exponent α of the velocity. For example, there are viscous dampers that have already been developed with variable values of α for wind and earthquake range of velocities, in particular α 5 1 (linear behavior) for low velocities related to wind and α ,0.1 (highly nonlinear behavior) for high velocity related to earthquake input (Infanti et al., 2008; Castellano et al., 2017). Further development will concern the change of the exponent α in more than two ranges, in order to optimize the dampers response to the different dynamic inputs, for example, frequent and rare wind and/or frequent and rare earthquakes. In alternative, the change of α could be done through semiactive devices as well.

9.2.2 NEW TECHNOLOGIES Continuing advances in material technologies are to be expected from research institutions and it is nearly certain that new stronger, lighter, and energyabsorbing materials will arrive. The present day technological exponential development also opens the gates for completely new solutions such as: •

Utilization of smarter materials, as classified by Ritter (2007): piezoelectric, electroactive, photostrictive, thermostrictive, magnetostrictive, chemostrictive materials, shape memory alloys, and fiber optic sensors. The major

9.2 Implementing Results of Technological Progress

•

•

applications at the moment are: cementitious materials, polymers nanocomposites, and ceramic nanocomposites. However, new materials are under development that could serve the scope of dissipating energy in the same manner as dissipation devices. Some of the most recent examples are: • Application of piezoelectric dampers that allow to control the motion based on the input without sensors and actuators. This solution was applied to a 30-story building in Japan (Marshall and Knapp, 2003). • Shape memory alloys have also been utilized as damping devices thank to their high dissipation capacity (Leong, 2005). • Nanomaterials have been utilized in the construction industry just recently (Silvestre, 2015). Development of a fluid with lower compressibility than silicone fluid. The latter is considered the best fluid for fluid viscous dampers, due to its high stability with temperature changes and high durability. However, its relatively high compressibility could be a problem in dampers for both wind and earthquake input, in which the wind displacement is much lower than the earthquake displacement and thus the compressibility reduces the energy dissipated per cycle. Harvest the energy that is dissipated by the damping devices to reduce the building vibrations and recycle the energy back to the energy supply system. The main idea is to generate the electricity by converting the kinetic energy (the rotation or swings of the pendulum or the relative displacement of the discrete damping elements) into voltage. One challenge for the harvesting system is the secure feed of the obtained energy back into the supply system or, alternatively, the use of energy storing devices. This could be interesting especially in the case of wind action. An example of such damping approach is the application of eddy current damper (ECD) systems (Fig. 9.1) (Lu et al., 2016). An eddy current is caused when a moving conductor intersects a stationary magnetic field, or vice versa. The relative motion between the conductor and the magnetic field generates the circulation of the eddy current within the conductor. These circulating eddy currents induce their own magnetic field with the opposite polarity of the applied field that causes a resistive force. These currents dissipate due to the electrical resistance and this force will eventually disappear. Hence, the

FIGURE 9.1 Scheme of an eddy current damping system.

927

928

CHAPTER 9 Future of dynamic modification systems

FIGURE 9.2 Scheme of a mechanical model of an inerter (Smith, 2002).

•

•

energy of the dynamic system will be removed. Since the resistive force induced by eddy currents is proportional to the relative velocity, the conductor and the magnet can be allowed to function as a form of viscous damping. ECD systems are contact-free devices which allow a maintenance-free design but they still require guides to ensure that the clearance between magnets and conductors remains constant for the design stroke of the damper system. This increases the design effort. The ECD systems also generate heat, so heat dissipation has to be addressed. For standard applications, the ECD systems are less cost-effective than conventional systems. The above-described ECD systems have a big potential to replace conventional damper systems for passive TMD systems since they allow larger strokes without having a significant increase in the damping media. Tuned mass inerter systems. The inerter is a linear device with two terminals free to move independently and it develops an internal (resisting) force proportional to the relative acceleration of its terminals (Fig. 9.2) (Smith, 2002). The constant of proportionality is called inertance and it causes a mass amplification, so less reaction mass as for a passive TMD system is required to achieve the required effectiveness. Despite his great potentiality, tuned mass inerter systems have not yet found their way as a practical application for larger structures. Fluid structure coupling (FSC) could be a new damping solution in the near future. This technology, developed by NASA’s Marshall Space Flight Centre in Huntsville (AL), is fully passive and highly efficient, that is, it only requires a fraction of the mass that is typically required by more traditional dampers (Morring, 2013). By engaging and mobilizing existing fluids within the structure (e.g., the water contained in a rooftop tank or in a swimming pool), this technology is able to control the way a tall building, for example, mitigates seismic or wind-induced motion.

CHAPTER

Tall building with dynamic modification systems trend data

10

CHAPTER OUTLINE 10.1 Database 1 (Worldwide Buildings Over 250 m) ................................................930 10.1.1 General Trends for Tall Buildings..................................................930 10.1.2 General Trends for Tall Buildings with Dynamic Modification Systems ..930 10.1.3 Trends for Dynamic Modification System Category.........................937 10.1.4 Trends for Dynamic Modification System Types .............................941 10.2 Database 2 (US Buildings Over 200 m)............................................................943 10.2.1 General Trends for Tall Buildings..................................................943 10.2.2 General Trends for Tall Buildings with Dynamic Modification System ......................................................................................945 10.2.3 Trends for Dynamic Modification System Category.........................947 10.2.4 Trends for Dynamic Modification System Type...............................948 10.3 Further Studies ..............................................................................................948 10.3.1 Dynamic Modification System Versus Height.................................949 10.3.2 Structural Material .....................................................................953 10.3.3 Structural System.......................................................................958 10.3.4 Building Function.......................................................................958 10.4 Summarized Data for Tall Buildings with Dynamic Modification System .............961 10.5 Conclusions...................................................................................................961

This chapter is devoted to study dynamic modification devices’ trends in highrise buildings constructed around the world. To this end buildings above an architectural height of 250 m are chosen for the analysis to include megatall (over 600 m) and supertall (over 300 m and below 600 m) buildings (based on CTBUH classification). In addition to the global trends, the study is further extended in the United States to a wider range of height, buildings above 200 m, in order to analyze dynamic modification devices’ trends somehow in one of the regions in which the use of supplementary damping is most-seen. The database source is the CTBUH skyscraper center and in the analysis it has been considered only those buildings that are complete, topped out (structurally, architecturally), and under construction by 2020. Excluded from the list are buildings on hold, never completed, and telecommunication/observation towers. For these reasons two different databases have been analyzed: •

Database 1: 597 buildings over 250 m located worldwide.

Damping Technologies for Tall Buildings. DOI: https://doi.org/10.1016/B978-0-12-815963-7.00010-5 Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved.

929

930

CHAPTER 10 Tall building with dynamic modification systems

•

Database 2: 229 buildings over 200 m located in the United States. Note that Database 2 also includes those US buildings part of Database 1.

To illustrate dynamic modification devices system trends, a comprehensive online searching is conducted for all the buildings included in these databases. Accordingly for those buildings that are equipped with a dynamic modification system, the essential data are collected (see Section 10.4 for a complete list), such as: • • • •

Device category (e.g., distributed, mass) Device type (e.g., viscous damper, TMD) Number and position of devices Main response application (controlling wind and/or earthquake)

It is important to note that for some buildings in both databases (39 buildings in Database 1 and 33 buildings in Database 2) no specific information about dynamic modification devices was found. In the following, at first, Database 1 is analyzed followed by the review of Database 2, using charts and maps. The main aspects that are studied in deep are the following: • • • • • • • •

General trends for tall buildings General trends for dynamic modification system Trends for dynamic modification system category Trends for dynamic modification system type Height Structural material Structural systems Building function

10.1 DATABASE 1 (WORLDWIDE BUILDINGS OVER 250 M) 10.1.1 GENERAL TRENDS FOR TALL BUILDINGS According to Fig. 10.1, among 597 buildings above 250 m in Database 1, 470 buildings (79%) are built in the recent decade (2010 2020), while there are 66 (11%) and 34 (6%) buildings built in last decades (2000 2010; 1990 2000, respectively). This demonstrates significant increasing trends for tall buildings over 250 m in the world in the recent decade. This shows the reason why tall buildings have gain so much attention in the last years, and also leading to an increasing demand for more efficient and environmental conscious solutions.

10.1.2 GENERAL TRENDS FOR TALL BUILDINGS WITH DYNAMIC MODIFICATION SYSTEMS This section describes general trends for supplementary damping systems in tall buildings available in Database 1. As seen in Fig. 10.2, among 559 buildings over 250 m (only those which dynamic modification system information could be found), 60 buildings (11%) are identified with supplementary damping (Fig. 10.3).

10.1 Database 1 (Worldwide Buildings over 250 m)

FIGURE 10.1 Trends for buildings over 250 m in the world by decade.

FIGURE 10.2 Dynamic modification system trends in tall buildings worldwide over 250 m.

Looking at the building trends with dynamic modification system in Fig. 10.3, it can be observed that the major use of supplementary damping in these buildings refers to the last three decades (about 97% of total). In particular, in the last decade tall buildings equipped with supplementary damping have increased more than two times compared with the previous decade; however, only 8% of the total building stock have been equipped with these devices (compared with 26% of the previous decade (2000 2010)). This trend can be better understood by looking at the building distribution in the world in relation with wind and seismic hazard maps as shown in Fig. 10.4. Looking into details of damping trends for different regions, Figs. 10.5 and 10.6 show that there are four main regions for tall buildings equipped with supplementary damping: Asia, North America, Middle East, and Australia. In general this may be related to the growing number of tall buildings in these regions as seen in Fig. 10.4. It is worth noticing that among 74 buildings in North America there are

931

932

CHAPTER 10 Tall building with dynamic modification systems

FIGURE 10.3 Dynamic modification system trends in buildings worldwide over 250 m by decade.

19 buildings with supplementary damping, meaning that 26% of the buildings are damped. While there are only 27 buildings with dampers among 346 buildings built in Asia, this is only 8% of total buildings. In the Middle East, only 11% of buildings over 250 m are with dynamic modification system. These results could show a higher application of damping systems in North America and possible reason of this could be: • • •

Considerable amount of damper manufactures (as shown in Chapter 4) Advanced standard codes for dynamic modification system design (as shown in Chapter 5) Dynamic modification system training in engineering school that leads to more qualified structural engineers

From the analyzes of the data, it has been found that most of the tall buildings with dynamic modification system are to be completed by 2020 as shown in Fig. 10.7. As shown in the charts, about 42% of damped buildings over 250 m in the world would be completed between 2017 and 2020, meaning that dynamic modification system trends in recent years have been increased significantly. In contrary, comparing the number of buildings with and without damper using Fig. 10.7, it can be seen that there is a slight decrease in damped building in the near future. Based on the dynamic modification system trends discussed earlier, it may be more desirable to focus mainly upon regions with more application of dynamic modification system as shown in Fig. 10.8, for the last three decades. It can be observed that the number of damped buildings is increasing in almost all regions, but the only region in which the percentage of damped buildings does not decrease in the recent decade is North America. This demonstrates the general conclusion drawn before that there is a general higher trend for tall damped buildings in North America.

FIGURE 10.4 Worldwide distribution of tall buildings (over 250 m).

FIGURE 10.5 Worldwide distribution of buildings with dynamic modification devices over 250 m by region.

10.1 Database 1 (Worldwide Buildings over 250 m)

FIGURE 10.6 Dynamic modification system trends in buildings over 250 m by region.

FIGURE 10.7 Dynamic modification system trends in buildings over 250 m in two time periods.

Given that the major quantity of damped buildings over 250 m were built or are to be completed in the recent decade, it could be interesting to analyze damping trends in countries of each region in the last decade, 2010 2020, as highlighted in Fig. 10.9. From the figure the following conclusions are drawn: •

Asia: only China, Japan, South Korea, Taiwan, and Philippines have damped buildings. China has the highest number but compared to the total building being constructed only 5% will be equipped with damping systems. Moreover in Japan and Taiwan all buildings over 250 m (in recent decade) have additional damping system.

935

FIGURE 10.8 Dynamic modification system trends in buildings over 250 m for major regions in last three decades.

10.1 Database 1 (Worldwide Buildings over 250 m)

FIGURE 10.9 Dynamic modification system trends in buildings over 250 m for major regions in recent decade (2010 2020).

•

•

North America: most of the damped buildings are located in the United States (with 12 buildings), while only three buildings have been built in Canada. However, this is 40% of the total constructed buildings in the United States and 66% of the one built in Canada. Middle East: United Arab Emirates (UAE) has the highest number of damped building with 3 out of 54 (i.e., 5% of total), while only one damped building has been constructed in Saudi Arabia.

These results reinstate again the major trend of damped tall building happening in North America compared to the rest of the world.

10.1.3 TRENDS FOR DYNAMIC MODIFICATION SYSTEM CATEGORY Given the general dynamic modification system trends discussed in previous sections, this section focuses only on tall buildings with damping system,

937

938

CHAPTER 10 Tall building with dynamic modification systems

FIGURE 10.10 Dynamic modification category trends in buildings over 250 m.

highlighting possible trends for dynamic modification system categorized as distributed, mass, active/semiactive/hybrid, and base isolation. Fig. 10.10 shows the damper category distribution for all the tall damped buildings in the world over 250 m (total of 60). It can be seen that the most utilized damper category is mass, that is, with 38 buildings (about 62% of total); while only 15 buildings (i.e., 25% of total) are equipped with distributed dampers and 7 with active (i.e., 12% of total). In the figure, it is also shown how there is a growth in utilizing distributed dampers in the last decade. Dominant tendencies for mass damper seem to be due to its higher efficiency in controlling wind-induced vibrations, as well as in satisfying occupant comfort criteria, which is the most prominent problems for supertall buildings (as described in Chapter 3). Instead, buildings equipped with distributed dampers are usually located in zones with significant seismicity or close to them (e.g., Chongqing and Xiamen in China, Los Angeles, and Seattle in the United States (Fig. 10.11)). This could be related to the higher effectiveness of distributed-type dampers than mass dampers in controlling seismic response (as discussed in previous chapters). Furthermore it is worth mentioning that no tall building above 250 m has been supported by base isolation. Looking on how the different damper categories are distributed in the different regions, Figs. 10.12 10.15 show a series of maps. As seen from Fig. 10.12,

FIGURE 10.11 Worldwide distribution of building with dynamic modification systems (over 250 m) by category.

940

CHAPTER 10 Tall building with dynamic modification systems

FIGURE 10.12 Dynamic modification category trends in buildings over 250 m in North America region between 2010 and 2020.

10.1 Database 1 (Worldwide Buildings over 250 m)

FIGURE 10.12 (Continued).

distributed-type dampers are mostly utilized in Asian countries, such as China, South Korea, and Philippines, with the highest applications in China. There is also two applications of such dampers in the United States. Major trends for mass-type damping system are in the United States, China, UAE, and Canada. Active-type damper is only utilized in Japan, Taiwan, China, and UAE.

10.1.4 TRENDS FOR DYNAMIC MODIFICATION SYSTEM TYPES In this section, dynamic modification-type tendencies in damped buildings are discussed. Fig. 10.16 shows that the distribution of the major damper types is viscous damper (82% of total) versus only 6% viscoelastic, 6% BRB, and 6% mixed (combination of more than one type of distributed devices). This trend toward viscous type could be related with the possibility to use these dampers for mitigating both wind and seismic response. Concentrating on the last two decades, Fig. 10.16 shows that distributed dampers have been widely utilized mainly in the last decade. With regard to general trends for mass dampers (Fig. 10.17), the major contribution is associated to TMD (66%), while the rest of mass-type dampers are

941

942

CHAPTER 10 Tall building with dynamic modification systems

FIGURE 10.13 Dynamic modification category trends in buildings over 250 m in Middle-East region between 2010 and 2020.

10.2 Database 2 (US Buildings over 200 m)

FIGURE 10.13 (Continued).

liquid devices (mainly TLD). To clearer understand tendencies of such damper types, Fig. 10.17 also shows that TLD are mainly utilized in the recent decade. To better understand what are the actual TMD and TLD trends in the major regions with damped buildings, Fig. 10.18 illustrates such trends in the last two decades. According to the charts of this figure, the number of buildings above 250 m equipped with TLD has grown in the recent decade in all the regions, especially in North America and Middle East. Instead, the quantity of TMDs utilized in the recent decade grows up in all the regions except in Middle East. Moreover no TMD/TLD was employed in Australia in the previous decade (2000 2010); however, few examples are seen in the recent decade.

10.2 DATABASE 2 (US BUILDINGS OVER 200 M) 10.2.1 GENERAL TRENDS FOR TALL BUILDINGS Database 2 includes 229 buildings over 200 m in the United States. Fig. 10.19 shows that the tall buildings industry started in the 1930s but the major development started in the 1970s. Subsequently there was a reduction in the last two

943

FIGURE 10.14 Dynamic modification category trends in buildings over 250 m in Asia region between 2010 and 2020.

10.2 Database 2 (US Buildings over 200 m)

FIGURE 10.14 (Continued).

decades due to different financial crises. However, in the current decade there is a great increase in the tall building industry with the construction of 76 buildings (about 25% of total). In the following, dynamic modification system trends for these buildings are reviewed, especially for those built in recent decades.

10.2.2 GENERAL TRENDS FOR TALL BUILDINGS WITH DYNAMIC MODIFICATION SYSTEM As seen from Fig. 10.20, among a total of 299 buildings over 200 m in the United States, only 196 have been found to have information about dynamic modification system. Among these only 35 buildings (18% of total) are found to be equipped with dynamic modification devices. Dynamic modification system trends by decade in the over 200-m US buildings are shown in Fig. 10.21. It is evident from the histogram that the major use of supplementary damping is in the current decade (69% of total). Moreover among 62 buildings of the recent decade, 24 buildings (39% of total) are supported by dynamic modification systems. This demonstrates increasing trends in the US buildings over 200 m in the recent decade. Moreover from Fig. 10.21 the

945

946

CHAPTER 10 Tall building with dynamic modification systems

FIGURE 10.15 Dynamic modification category trends in buildings over 250 m in South-East Asia and Oceania regions between 2010 and 2020.

10.2 Database 2 (US Buildings over 200 m)

FIGURE 10.15 (Continued).

pie charts show that among the building under completion by 2020 almost half of them (46%) will be equipped with dynamic modification systems. This proves once again the considerable trends for supplementary damping in recent years.

10.2.3 TRENDS FOR DYNAMIC MODIFICATION SYSTEM CATEGORY Tendencies for dynamic modification system category are shown in Fig. 10.22. The figure shows that only two dynamic modification categories are employed in the United States: 30 mass (88%) and 4 distributed (12%). Leading tendencies of mass dampers probably may refer to their main role in controlling wind response (occupant comfort criteria), since most of these buildings are located in zones with considerable wind storm hazard (and low-seismic hazard, Fig. 10.12). It is worth noticing that most of the distributed dampers are employed in buildings located near high seismic hazard regions (Los Angeles, San Francisco, and Seattle, Fig. 10.12). Also none of the active-type systems or base isolation is utilized. In addition, it can be seen from Fig. 10.22 that there are only two distributed dampers in the last two decades and all the rest are mass dampers.

947

948

CHAPTER 10 Tall building with dynamic modification systems

FIGURE 10.16 Distributed damper-type trends in buildings over 250 m.

10.2.4 TRENDS FOR DYNAMIC MODIFICATION SYSTEM TYPE As the majority of supplementary damping building in the United States were mass type, only trends for this type are discussed in this section. Fig. 10.23 shows that TMD is mostly employed (67% of total). To understand what trends govern the last two decades, Fig. 10.23 shows that TMD is the most utilized systems in both decades. It is important to note that even though a few TLCD is utilized in the current decade, no application of such a dynamic modification system is seen in the current decade.

10.3 FURTHER STUDIES In this section other parameters are analyzed, such as height and structural details, and building functions are investigated.

10.3 Further Studies

FIGURE 10.17 Mass damper-type trends in buildings over 250 m.

10.3.1 DYNAMIC MODIFICATION SYSTEM VERSUS HEIGHT It is important to understand if supplementary damping is correlated to the height of the buildings and in particular to the most iconic building being built in the world. Fig. 10.24 shows how 25% of the top 40 and 30% of the top 20 tall buildings are equipped with dynamic modification devices. This data compared with the full Database 1 (that has only 11% of building equipped with dynamic modification devices) show that as building get taller supplementary damping system could be a more viable solution. Table 10.1 shows the tallest 40 buildings (with and without damper) for Database 1 with a summary of the major data. As it can be seen that TMD is the most utilized dynamic modification system. As defined by CTBUH, buildings can be classified as megatall (over 600 m), supertall (between 300 and 600 m), and tall (between 50 and 300 m). These

949

950

CHAPTER 10 Tall building with dynamic modification systems

FIGURE 10.18 TMD and TLD trends in buildings over 250 m in last two decades by different regions.

criteria can be useful to see if there is any correlation between the building height and supplementary damping as shown in Fig. 10.25. It can be seen that in megatall building, only mass dampers are employed. This is due to dominant performance of such systems in satisfying occupant comfort in upper stories during wind storms. Therefore, distributed and active-based systems are only utilized in buildings under 600 m. Also in supertall and tall buildings, majority of dynamic modification systems are of mass type: 64% of total in supertall and 57% in tall buildings. Note that the tallest building in which mass damper is installed is Jeddah Tower (Jeddah, Saudi Arabia) with 1000 m height, while Wilshire Grand Center (Los Angeles, USA) with the height of 335 m is the tallest with distributed damper. The tallest building with active-based control systems, Shanghai World Financial Center (Shanghai, China), is 492 m high. Fig. 10.26 shows the

FIGURE 10.19 Trends for buildings over 200 m in the United States by decade.

952

CHAPTER 10 Tall building with dynamic modification systems

FIGURE 10.20 Dynamic modification system trends in tall buildings over 200 m in the United States.

FIGURE 10.21 Dynamic modification system trends in US buildings over 200 m by time.

10.3 Further Studies

FIGURE 10.22 Dynamic modification category trends in US buildings over 200 m.

worldwide distribution of the first top 20 tallest building with dynamic modification system in the world.

10.3.2 STRUCTURAL MATERIAL According to Table 10.1, Fig. 10.27 analyzes how dynamic modification systems related to different structural materials (steel, concrete, composite, and mixed structure). As stated by the CTBUH criteria, “Composite is a combination of both steel and concrete components used together in the main structural elements”; and “Mixed-Structure utilizes distinct steel and concrete systems, one on top of the other. Steel/concrete indicates a steel structural system located on top of a concrete structural system, with the opposite concrete/steel.” As shown in Fig. 10.27, most of the buildings with damper are in concrete and composite, meaning that the main structural systems are in concrete material and only 11% of damped buildings are in steel.

953

954

CHAPTER 10 Tall building with dynamic modification systems

FIGURE 10.23 Mass damping-type trends in US buildings over 200 m.

FIGURE 10.24 Dynamic modification trends in tallest buildings.

Table 10.1 List of 40 Tallest Building According to Skyscraper Center Number

Building

City

Country

Height (m)

Floors

Structural Material

Main Lateral Structural System

Damper Category

Damper Type

1

Jeddah Tower

Jeddah

Saudi Arabia

1000

167

Composite

Yes

Isolated

2

Burj Khalifa

Dubai

828

163

Steel/concrete

3

Wuhan Greenland Center Shanghai Tower Makkah Royal Clock Tower Ping An Finance Center Goldin Finance 117 Global Financial Center Tower 1 Lotte World Tower One World Trade Center Guangzhou CTF Finance Centre Tianjin CTF Finance Centre China Zun Tower

Wuhan

United Arab Emirates China

636

125

Composite

Residential, serviced apartments, hotel Hotel, residential, office Hotel, residential, office

Shanghai

China

632

128

Composite

Hotel, office

Yes

Mecca

Saudi Arabia

601

120

Steel/concrete

Other, hotel

No

Shenzhen

China

599

115

Composite

Office

Yes

Tianjin

China

596

128

Composite

Hotel, office

No

Shenyang

China

568

114

Concrete

Office

No

Seoul

South Korea

554

123

Composite

Hotel, office

No

New York City Guangzhou

United States

541

94

Composite

Office

No

China

530

111

Composite

Hotel, residential, office

No

Tianjin

China

530

97

Composite

No

Beijing

China

528

108

Composite

Hotel, serviced apartments, office Office

4 5 6 7 8 9 10 11

12 13

No No

Isolated

Isolated

No (Continued)

Table 10.1 List of 40 Tallest Building According to Skyscraper Center Continued Number

Building

City

Country

Height (m)

Floors

Structural Material

Main Lateral Structural System

Damper Category

14

Entisar Tower

Dubai

520

111

Concrete

Dalian Greenland Center TAIPEI 101 Shanghai World Financial Center International Commerce Centre Central Park Tower Corporate Avenue 1 Lakhta Center

Dalian

518

88

Composite

Residential, serviced apartments, hotel Hotel, residential, office

No

15

United Arab Emirates China

Taipei Shanghai

Taiwan China

508 492

101 101

Composite Composite

Office Hotel, office

Yes Yes

Hong Kong

China

484

108

Composite

Hotel, office

No

New York City Chongqing

United States

472

95

Concrete

No

China

468

99

Composite

Residential, hotel, retail Hotel, office

No

Russia

462

86

Composite

Office

No

Vietnam

461

81

Composite

No

China

452

94

Composite

SOHO, hotel, residential Hotel, office

No

Malaysia

452

88

Composite

Office

Yes

Isolated

Malaysia

452

88

Composite

Office

Yes

Isolated

China

450

98

Composite

Yes

Isolated

China United States

450 435

66 82

Composite Concrete

Hotel, office, serviced apartments Hotel, office Residential

No Yes

Isolated

16 17 18

19 20 21 22 23 24 25 26 27 28

Vincom Landmark 81 Changsha IFS Tower T1 Petronas Twin Tower 1 Petronas Twin Tower 2 Suzhou IFS Zifeng Tower 111 West 57th Street

St. Petersburg Ho Chi Minh City Changssha Kuala Lumpur Kuala Lumpur Suzhou Nanjing New York City

Damper Type

No

Isolated Active

29

30 31 32 33 34

35 36 37 38

39 40

Nanning China Resources Tower Marina 106 Willis Tower World One KK100 Guangzhou International Finance Center Wuhan Center Tower Riverview Plaza A1 432 Park Avenue Multifunctional Highrise Complex Akhmat Tower Diamond Tower Haikou Tower 1

Nanning

China

445

94

Composite

Hotel, office

No

Dubai

445

104

Concrete

Residential

No

Chicago Mumbai Shenzhen Guangzhou

United Arab Emirates United States India China China

442 442 442 439

108 117 100 103

Steel Composite Composite Composite

Office Residential Hotel, office Hotel, office

No No No No

Wuhan

China

438

88

Composite

No

Wuhan

China

436

73

Composite

Hotel, residential, office Hotel, office

No

New York City Groznyj

United States

426

85

Concrete

Residential

Yes

Russia

435

102

Steel

Residential, office, hotel

No

Jeddah Haikou

Saudi Arabia China

432 428

93 94

Composite

Residential Hotel, residential, office

No No

Isolated

958

CHAPTER 10 Tall building with dynamic modification systems

FIGURE 10.25 Height criteria for dynamic modification categories.

10.3.3 STRUCTURAL SYSTEM To study the correlation between dynamic modification and structural systems, the following tall building structural systems are considered, as classified by Gu¨nel and Ilgin (2014): • • • • • •

Frame systems (rigid; braced) Core systems (RC shear wall systems) Shear-frame systems (shear trussed frame/braced frame systems; shear walled frame systems) Mega frame (mega column; space truss) Outrigger systems (core supported to perimeter columns) Tube systems (framed-tube systems; trussed-tube systems; bundled-tube systems)

As shown in Fig. 10.28, most of the buildings with damper are reinforced by outrigger systems representing 43% of total. This demonstrates that even though stiffening the structural system using outriggers efficiently resists the building against lateral loadings, but supplementary damping was required for response control and occupant comfort criteria. Other structural systems such as tube systems, core systems, and shear-frame systems can be seen among the structural system employed in damped buildings.

10.3.4 BUILDING FUNCTION Function of buildings with damper is analyzed based on the CTBUH definitions (Fig. 10.29) consisting of residential, hotel, office, and mixed-use. As per CTBUH criteria, “a mixed-use tall building contains two or more functions, where each of the functions occupies a significant proportion of the tower’s total space.” As seen from

FIGURE 10.26 World distribution of top 20 tallest buildings with dynamic modification systems.

960

CHAPTER 10 Tall building with dynamic modification systems

FIGURE 10.27 Structural materials used in buildings with dynamic modification system.

FIGURE 10.28 Structural systems used in buildings with dynamic modification system.

FIGURE 10.29 Structural systems used in buildings with dynamic modification system.

10.5 Conclusions

Fig. 10.29, most of the damped buildings have a mixed-use function (as a combination of office, residential, and/or hotel).

10.4 SUMMARIZED DATA FOR TALL BUILDINGS WITH DYNAMIC MODIFICATION SYSTEM Tables 10.2 and 10.3 summarize the most useful data of the buildings with dynamic modification system between Databases 1 and 2, respectively. They include the building name, city, country, height, number of floors, damper category and its type, and structural systems and materials. Note that buildings have been listed in ascending height order.

10.5 CONCLUSIONS The main conclusions obtained from the evaluation of dynamic modification system trends in tall buildings (for the specified height limit) are summarized as follows: •

• • • • • • • • • • • •

Most of the buildings analyzed are not equipped with additional damping. This may denote that dynamic modification system is not the only strategy to control dynamic response. Instead, other remedies such as adding stiffness or modifying aerodynamic shape of buildings could be utilized. However, among the top 20 buildings 30% of the building have been equipped with dynamic modification systems. Major regional trends for dynamic modification systems are in China and North America. Utilization of dynamic modification devices in tall buildings is rising in the next upcoming years, 2017 2020. Most dynamic modification system trends, especially in the recent decade, are related to North America, especially United States. Although most of the supertall buildings have been built in China, a small part of them is equipped with dynamic modification system. Most of the supplementary damping employed is of mass type, mostly TMD. TLD utilization have increased considerably in the recent decade. Trends for distributed dampers are increasing in the recent decade, mainly in Asia, with majority in China. The dominant part of distributed dampers is of viscous type versus a few displacement-dependent devices. A few utilization of active-based dampers is seen, mostly in Japan. No base isolation record is found in buildings over 250 m. In megatall buildings (over 600 m), only mass dampers (TMD or TLD) are employed. Despite the use of efficient structural systems such as tube and outrigger systems in many tall buildings, additional damping is sometimes required for response control.

961

Table 10.2 List of Tall Buildings With Dynamic Modification System Worldwide up to 250 m Building

City

Country

Height (m)

Floors

Structural Material

Main Lateral Structural System

Damper Category

Damper Type

Number 1 2 3

Jeddah Shanghai Shenzhen

Saudi Arabia China China

1000 1 632 599

167 128 115

Composite Composite Composite

Shear-frame system Outrigger system Outrigger system

Mass Mass Mass

TMD TMD TMD

Taipei Shanghai

Taiwan China

508 492

101 101

Composite Composite

Outrigger system Outrigger system

Mass Active

TMD ATMD

Tianjin

China

468

91

Composite

Distributed

Viscous

Kuala Lumpur

Malaysia

452

88

Composite

Outrigger system

Mass

TMD

8 9 10 11

Jeddah Tower Shanghai Tower Ping An Finance Center Taipei 101 Shanghai World Financial Center R&F Guangdong Building Petronas Twin Tower 2 Suzhou IFS 111 West 57th Street 432 Park Avenue Princess Tower

Suzhou New York City New York City Dubai

450 435 426 413

98 82 85 101

Mass Mass Mass Mass

TLD TMD TMD TLD

23 Marina

Dubai

392

88

Composite Concrete Concrete Steel/ concrete Concrete

Outrigger system Outrigger system Shear-frame system Shear-frame system

12

Outrigger system

Mass

TMD

13

30 Hudson Yards

New York City

China United States United States United Arab Emirates United Arab Emirates United States

387

73

Core system

Mass

TMD

14 15

Vista Tower Almas Tower

Chicago Dubai

363 360

101 68

Core system Tube system

Mass Mass

TLD TMD

16

Emirates Tower One

Dubai

355

54

Composite

Tube system

Mass

TMD

17

Raffles City Chongqing T3N

Chongqing

United States United Arab Emirates United Arab Emirates China

Steel/ concrete Concrete Concrete

354

79

Concrete/ steel

Outrigger system

DistributedIsolation

18

Raffles City Chongqing T4N

Chongqing

China

354

79

Concrete/ steel

Outrigger system

DistributedIsolation

Viscous damper—LRB, FPB Viscous damper—LRB, FPB

4 5 6 7

19 20

Kaohsiung Los Angeles

Taiwan United States

348 335

85 73

Composite Composite

Mega frame outrigger system

Active Distributed

HMD BRB

Shanghai

China

333

60

321

56

Mass

TMD

23

53 West 53rd

New York City

United Arab Emirates United States

Shear-frame system with outrigger Mega frame

Viscous damper

Dubai

Concrete/ steel Composite

Distributed

22

T & C Tower Wilshire Grand Center Shimao International Plaza Burj Al Arab

320

77

TMD

Australia 108

Melbourne

Australia

317

100

Mass

TLD

25

Emirates Tower Two

Dubai

309

56

Composite

Mass

TMD

26 27 28

New York City New York City Incheon

308 306 305

72 75 68

Concrete Concrete Concrete

Mass Mass Distributed

TMD TLD Viscous damper

29

35 Hudson Yards One57 Northeast Asia Trade Tower Abeno Harukas

United Arab Emirates United States United States South Korea

Diagrid-outrigger systems Core and outrigger systems Tube systems

Mass

24

Concrete/ steel Concrete

Osaka

Japan

300

60

Composite

30

Aspire Tower

Doha

Qatar

300

36

Composite

Active/ distributed Mass

ATMD viscousfriction TMD

31

Eureka Tower

Melbourne

Australia

297

91

Concrete

32 33 34

Comcast Center Landmark Tower Columbia Center

Philadelphia Yokohama Seattle

United States Japan United States

297 296 291

57 73 76

35

220 Central Park South

New York City

United States

289

66

21

Mass

TLCD

Concrete Steel Composite

Core systems Core 1 outrigger frame systems Core 1 outrigger systems Core systems 1 steel mesh Outer tube 1 outrigger 1 core Core Tube systems Brace moment frame

Mass Hybrid Distributed

Concrete

Shear-frame systems

Mass

TLCD HMD Viscoelastic damper TMD (Continued)

Table 10.2 List of Tall Buildings With Dynamic Modification System Worldwide up to 250 m Continued Building

City

Country

Height (m)

Floors

Structural Material

Main Lateral Structural System

Damper Category

Damper Type

Number 36

Citicorp Building 601 Lexington

New York City

United States

279

63

Steel

Frame systems 1 supertruss

Mass

37

125 Greenwich Street Nan Shan Plaza Aura at College Park Crown Sydney Hotel and Resort 21st Century Tower

New York City

United States

278

72

Concrete

-

Mass

TMD with hydraulic actuator TMD

Taipei Toronto Sydney

Taiwan Canada Australia

272 272 271

48 78 75

Frame systems Shear-frame systems core systems

Mass Mass Mass

TMD TLD TMD

Dubai

269

55

Shear-frame systems

Mass

TMD

265

56

Steel Concrete Concrete/ steel Concrete/ steel Concrete/ steel

Outrigger system

DistributedIsolation

38 39 40 41 42

Raffles City Chongqing T2

Chongqing

United Arab Emirates China

43

Raffles City Chongqing T5

Chongqing

China

265

56

Concrete/ steel

Outrigger system

DistributedIsolation

44

Raffles City Chongqing T3S

Chongqing

China

265

46

Concrete/ steel

Outrigger system

DistributedIsolation

45

Raffles City Chongqing T4S

Chongqing

China

265

50

Concrete/ steel

Outrigger system

DistributedIsolation

46 47 48

Triumph Palace Trump World Tower The Address Residence Sky View Tower 1 Grand Hyatt Metrocenter

Moscow New York City Dubai

Russia United States United Arab Emirates

264 262 261

61 72 60

Concrete Concrete Concrete

Tube systems

Mass Mass Mass

Viscous damper—LRB, FPB Viscous damper—LRB, FPB Viscous damper—LRB, FPB Viscous damper—LRB, FPB TMD TMD TMD

Taguig City

Philippines

258

66

Concrete

Outrigger system

Distributed

Viscous damper

49

50 51 52 53 54

Greenland Center 1 Greenland Center 2 Hyatt Park Tower One Bennett Park Rinku Gate Tower

Urumqi Urumqi Chicago Chicago Izumisano

China China United States United States Japan

258 258 257 257 256

57 57 68 77 56

55

Osaka World Trade Center Toranomon Hills

Osaka

Japan

256

Tokyo

Japan

Toronto Hong Kong New York City Beijing

56 57 58 59 60

One Bloor Street East Highcliff Apartments 56 Leonard Street Beijing Yintai Center

Outrigger system Outrigger system Shear-frame system Shear-frame system

Distributed Distributed Mass Mass Hybrid

Viscous damper Viscous damper TMD TLD HMD

55

Composite Composite Concrete Concrete Concrete/ steel Steel

Frame system

Hybrid

HMD

256

52

Steel

Tube system

Distributed

Canada

255

75

Concrete

Core system

Mass

Oil, BRB, Friction TLD

China United States China

252 250 250

73 57 62

Concrete Concrete Concrete

Shear-frame system Outrigger system Shear-frame system

Mass Mass Distributed

TLD TLD Viscous damper

Table 10.3 List of Tall Buildings With Dynamic Modification System in USA up to 200 m Building

City

Country

Height (m)

Floors

Structural Material

Main Lateral Structural System

Damper Category

Damper Type

Number 1

111 West 57th Street

United States

435

82

Concrete

Outrigger system

Mass

TMD

2

432 Park Avenue

United States

426

85

Concrete

Shear-frame system

Mass

TMD

3

30 Hudson Yards

United States

387

73

Mass

TMD

Vista Tower Wilshire Grand Center

United States United States

363 335

101 73

Steel/ concrete Concrete Composite

Core system

4 5

Core system Outrigger system

Mass Distributed

TLD BRB

6

53 West 53rd

United States

320

77

Mass

TMD

35 Hudson Yards

United States

308

72

Concrete/ steel Concrete

Diagrid-outrigger systems

7

Mass

TMD

8

One57

United States

306

75

Concrete

Core systems

Mass

TLD

9 10

Comcast Center 220 Central Park South

United States United States

297 291

57 66

Concrete Concrete

Core

Mass Mass

TLCD TMD

11

Columbia Center

New York City New York City New York City Chicago Los Angeles New York City New York City New York City Philadelphia New York City Seattle

United States

291

76

Composite

Brace moment frame

Distributed

12

220 Central Park South

United States

289

66

Concrete

Shear-frame systems

Mass

Viscoelastic damper TMD

13

Citicorp Building Lexington

United States

279

63

Steel

Frame systems 1 supertruss

Mass

14

125 Greenwich Street

United States

278

72

Concrete

15

Trump World Tower

United States

262

72

Concrete

16 17

Hyatt Park Tower One Bennett Park

United States United States

257 257

68 77

Concrete Concrete

601

New York City New York City New York City New York City Chicago Chicago

Mass

TMD with hydraulic actuator TMD

Tube systems

Mass

TMD

Shear-frame system Shear-frame system

Mass Mass

TMD TLD

18

56 Leonard Street

19

Bloomberg Tower

20

181 Fremont

21

111 Murray Street

22 23

200 Clarendon John Hancock Tower 520 Park Avenue

24

50 West

25

Madison Square Park Tower Goldman Sachs Headquarters Two Union Square

26 27 28 29 30

FMC Tower at Cira Centre South 150 North Riverside Metropolitan Tower

31

100 East 53rd Street

32 33 34

The Independent Waldorf Astoria Chicago Random House Tower

35

The Austonian

New York City New York City San Francisco New York City Boston

United States

250

57

Concrete

Outrigger system

Mass

TLD

United States

246

55

Outrigger system

Mass

TMD

United States

244

54

Concrete/ steel Steel

Mega frame

Distributed

Viscous damper

United States

241

58

Concrete

Shear-frame system

Mass

TMD

United States

241

62

Steel

Frame system

Mass

TMD

New York City New York City New York City New York City Seattle

United States

237

54

Concrete

Shear-frame system

Mass

TMD

United States

237

64

Concrete

Shear-frame system

Mass

TMD

United States

236

61

Concrete

Outrigger system

Mass

TMD

United States

228

44

Steel

Frame system

Mass

TMD

United States

226

56

Composite

Frame system

Distributed

Philadelphia

United States

222

49

Composite

Shear-frame system

Mass

Viscoelastic damper TLD

Chicago New York City New York City Austin Chicago

United States United States

221 218

53 68

Composite Concrete

Core system Shear-frame system

Mass Mass

MTMD TMD

United States

217

61

Concrete

Outrigger system

Mass

TMD

United States United States

209 209

62 60

Concrete Concrete

Outrigger system

Mass Mass

TLD TLD

New York City Austin

United States

208

52

Steel

Frame system

Mass

TLCD

United States

208

56

Concrete

Shear-frame system

Mass

TMD

CHAPTER

Conclusions

11

The results of this book were a 3-year effort conducted from the Council on Tall Building and Urban Habitat research group at the University IUAV, Venice. Different experts in the field (from academia, engineering consultants, developers) have contributed to the different chapters of the book. The combination of person coming from different building sectors helped to get different point of views on the importance and the utilization of dynamic modification device technologies for tall buildings. This document is a sort of manual on the state of the art for the application of dynamic modification devices for tall buildings. The scope was to provide the reader with the most up-to-date and comprehensive theory and practice on the subject. Given the large scope of the publication, it is important to understand the possible limitations in some of the arguments discussed. Indeed, the review mainly focused on the available literature written in English; therefore some of the codes and article not available in English could not be reviewed in detail. However, it is important to understand that the general goal was to provide a deep and broad overview on the different possibilities that dynamic modification devices can lead in the design of tall buildings. Both advantages and disadvantages in the utilization of this technology were discussed to provide an unbiased overview on the topic. A deep review of the different devices available has shown the great importance that damping system has gained in the building community (especially in the last 30 years). Indeed, given the great amount of research and experimental tests conducted, the reliability of these devices is really high even more than other building materials. Moreover, national building codes require manufactures to undergo stringent quality controls and tests to verify the behavior of the different devices. Most of the time engineers are not trained to design damping system; this is why the scope of this book was also to provide a general overview and some simple guidelines for the design of the different devices that are available in the market. In addition, the presented case studies have shown the prominent and important steps beyond planning, design, construction, and maintenance of buildings with added damping devices. The editors and authors hope that this document could become an important reference in this area of the tall building industry. Moreover, the deep review of the state of the art could be a good start for further research investigation in the development of this field as stated in the possible future development chapter.

Damping Technologies for Tall Buildings. DOI: https://doi.org/10.1016/B978-0-12-815963-7.00011-7 Copyright © 2019 Council on Tall Buildings and Urban Habitat (CTBUH). Published by Elsevier Inc. All rights reserved.

969

APPENDIX

A The general theories behind the major devices developed in Chapter 4 are reviewed here with relevant mechanical formulations.

A.1 FLUID VISCOUS DAMPERS The theory for both linear and nonlinear viscous damper types is reviewed here in detail. The discussion will refer to an SDOF system, and the same formulations will be extended to an arbitrary MDOF system including viscous dampers. In the case of LVDs, the force
velocity relationship is linear; from Equation (4.4) the following expression can be derived: Fdv 5 cv u_ d

(A.1)

The equivalent (linear) viscous damping (ζ eq ) of the damper can be calculated by Equation (3.60). In this case, the energy dissipated (ED ) under a full cycle of harmonic motion at resonance condition ud 5 u0 sinωt can be given as follows (Fig. A.1): þ ED 5

Fd dud 5

ð 2π=ω

Fdv u_ d dt 5

ð 2π=ω

0

0

cv u_ d 2 dt 5 πcv ωu20

(A.2)

where ω is the frequency of harmonic motion (frequency of SDOF system). The maximum strain energy of the structure is given by Equation (3.61). Substituting Equations (A.2) and (3.61) into Equation (3.60), it is possible to compute the equivalent viscous damping at ω as follows: ζ eq 5

Tcv 4πm

(A.3)

where T and m are the period and mass of the SDOF system, respectively. Inherent to the equations above, the resulting cv value is in units of force
time/ displacement (when the mass is divided by the acceleration of gravity) and a linear relationship between damper force and velocity is implied. Therefore, quantifying a damping level in terms of percent critical ccr (Equation (3.7)) becomes difficult when using a nonlinear damper.

971

972

APPENDIX A

FIGURE A.1 Hysteresis loop for fluid viscous dampers with three values of α (undergoing harmonic motion).

Equation (A.3) can be simply extended to an MDOF system as follows: ζ eq;m 5

P

2 j Nj cv;j φrj;m P 2 4π i mi φim

Tm

(A.4)

where cv;j is the (linear) viscous damper constant at j-th storey; Nj is the number of identical dampers with the same cv;j ; mi is the concentrated mass of level (floor) i; φrj;m is the relative displacement between two damper ends at j-th floor, corresponding to m-th mode; and φim is m-th mode displacement at level (floor) i. From Equations (A.3) and (A.4), it is obvious that the equivalent viscous damping is independent of the input demand, since it depends only on the structural system properties. This is not be the case for nonlinear viscous and viscoelastic dampers as shown in the following sections. In the case of an NLVD, if a comparison is made between the sinusoidal inputs of a fluid damper of the energy dissipated for each cycle to the damping exponent, a fluid damper with a lower damping exponent dissipates more energy per cycle than one with a higher exponent. Fig. A.2 shows a comparison between different exponent values. For α 5 0:4, it is demonstrated that the area under the force versus displacement curve (hysteresis) is higher than the response with α 5 1 (i.e., linear damper). However, for damping exponent less than 0.4, the increase in energy dissipation is minor as shown in Fig. A.2. In case if an NLVD equivalent (linear) viscous damping is needed, it cannot be derived exactly since it depends on the excitation demand on the system.

A.1 Fluid Viscous Dampers

FIGURE A.2 Comparison of energy absorbed with varying damping exponents with sinusoidal input.

Therefore, several “approximated” expressions were derived. The four most cited ones are as follows: •

•

•

•

Soong and Constantinou (1994), updated in Seleemah and Constantinou (1997) (called “SC” from now on). The equivalent (linear) damping coefficient expression is based on the assumption that the work done by an NLVD, in one cycle of vibration, can be equated to the energy dissipated by a linear viscous damping system (energy balance). Silvestri et al. (2010) simplified the Christopoulos and Filiatrault (2006) expression (called “SIL” from now on). The proposed expression is a further simplified version of the one proposed by Soong and Constantinou (1994). Lin et al. (2003) (called “LIN” from now on). The equivalent (linear) damping coefficient expression is based on the assumption that the energy dissipated by LVD and NLVD, in an SDOF system subjected to all cycles of harmonic motion, is the same. Pekcan et al. (1999) (called “PEK” from now on). The equivalent (linear) damping coefficient expression is based on the assumption that the timeaveraged power consumption over one cycle of sinusoidal loading can be approximated as the area under the force
velocity response curve (i.e., the rate of energy dissipated rather than energy as the other two methods).

The different formulations have a similar format that can be expressed as follows (Lago, 2011): ζ eq 5 κ

cv uα21 T 22α 0 ð2πÞ22α m

(A.5)

973

974

APPENDIX A

where κ is a coefficient that depends on the velocity power factor α. This factor can be computed as suggested by different authors as follows: ψ 2π

(A.6)

κSIL 5

1 2U0:8α21

(A.7)

κLIN 5

3ψ 2πð2 1 αÞ

(A.8)

κSC 5

1 ðα 1 1Þ   221α Γ2 1 1 α=2 ψ5 πΓð2 1 αÞ κPEK 5

(A.9) (A.10)

where ΓðÞ is the gamma function. In Equation (A.5), the equivalent viscous coefficient is dependent on the demand of the structure (uα21 ). Moreover, the equation 0 is exact only under sinusoidal excitation with a maximum displacement of u0 (often be set equal to 0.2
0.3 times maximum displacement of the damper). This is particularly critical for seismic excitations since in every cycle of vibration the equal energy dissipation between NLVD and equivalent linear viscous system is not guaranteed. The κ has been computed for different values of α in Fig. A.3 (normalized to the maximum value at α 5 1) by Lago (2011). As can be seen from Fig. A.3, there are two major trends: the SC and SIL expressions are very similar (since the SIL expression is an approximation of the SC), while the LIN and PEK formulations show a similar trend with a small difference (,1%) that increases as α decreases. Therefore, the first two expressions predict smaller damping values (difference of 17% at 0.5 and of 28% at 0.2) and the curves converge as α

FIGURE A.3 Equivalent damping formulation comparison. Adapted from Lago, A., 2011. Seismic Design of Structures with Passive Energy Dissipation Systems. IUSS Report, Pavia, Italy.

A.2 Viscoelastic Dampers

approaches 1. Comparing the results in terms of the dampers’ constant, given the same equivalent viscous damping constant, the first two expressions give higher values of cv . A study was conducted by Lago (2011) on the different expressions for an SDOF equipped with a damper. A series of time-history analyses were carried out based on 20 records from 10 events, for a range of equivalent damping values and velocity exponents. The results show that the expressions by Peckan et al. (1999) and Lin et al. (2008) provide better response compared to the expression by Soong and Constantinou (1994), which produces, in general, lower displacement values. In the case of MDOF structures equipped with dampers of similar α value, Equation (A.5) for the m-th mode can be defined as follows: P Tm 22α ζ eq;m 5 κ

j 2π

P

Aα21 Nj cv;j φ11α rj;m

i

mi φ2im

(A.11)

where cv;j is the nonlinear viscous damper constant for level (floor) j; κ is the factor as defined from Equations (A.6) to (A.9); A is the roof response amplitude related to modal displacement shape, φi , normalized to unit value at the roof. An important consideration of equivalent viscous damping is that the damping is assumed to be distributed throughout the building structure. In modern structures often times, for architectural and economic reasons, dampers are concentrated in a few locations. In such cases, it is recommended to utilize other numerical techniques whereby dampers are explicitly modeled in the analysis, such as free vibration, to determine the added modal damping provided by the dampers.

A.1.1 TEMPERATURE DEPENDENCY Constantinou et al. (1993) propose a study on a three-story test structure equipped with viscous dampers. The dynamic tests show that the period of the structure is essentially not affected by the insertion of fluid dampers. However, higher modes are stiffened since the frequencies are greater than the cutoff frequency of 4 Hz (Christopoulos and Filiatrault, 2006). In this study, a variation in the damping properties with temperature was tested and it was found that the damping constant is not greatly affected between temperatures of 24 C and 50 C (Fig. A.4).

A.2 VISCOELASTIC DAMPERS In the following text, the theory viscoelastic damper is reviewed in detail based on considerations given in Chapter 4.

975

APPENDIX A

10

Temp. = 0°C CL = 22.3 N.s/mm

8 Peak force (kN)

976

Temp. = 24°C CL = 15.5 N.s/mm Temp. = 50°C CL = 11.6 N.s/mm f < 5 Hz f > 5 Hz

6 4 2 0 0

100

200

300

400

500

Peak velocity (mm/s)

FIGURE A.4 Evaluation of damping constant for fluid damper (Constantinou et al., 1993).

Based on the viscoelastic model, a simple Kelvin solid model can be utilized as shown in Fig. 4.27. The shear stress
strain relationship at time, t, of a viscoelastic solid can be expressed as τ S ðtÞ and γ S ðtÞ, respectively: τ E ðtÞ 1 τ C ðtÞ 5 τ S ðtÞ

(A.12)

and by shear strain compatibility: γ E ðt Þ 5 γ C ðt Þ 5 γ S ðt Þ

(A.13)

where τ E ðtÞ and γ E ðtÞ are the elastic shear stress and strain at time t, respectively, and τ C ðtÞ and γ C ðtÞ are the viscous shear stress and strain of the material at time t, respectively. The elastic and viscous shear stress can be expressed as functions of the elastic shear modulus, GE , and a shear viscous damping constant, GC , as: 

τ E ðt Þ 5 G E γ S ðt Þ τ C ðtÞ 5 GC γ_ S ðtÞ

(A.14)

Then substituting Equation (A.14) with Equation (A.12), the shear constitutive relationship for the Kelvin solid model can be expressed as follows: τ S ðtÞ 5 GE γ S ðtÞ 1 GC γ_ S ðtÞ

(A.15)

The shear storage modulus, GE , can be estimated from several empirical formulas presented in the literature for different VE materials (Chang et al., 1991, 1995; Abbas and Kelly ,1993). Considering the force
displacement relationship shown in Equation (4.6), under a cycle of harmonic motion at resonance condition, the axial force in the viscoelastic dampers can be expressed as: Fd;ve ðtÞ 5 kve u0 sinωt 1 cve u0 ωcosωt

(A.16)

A.2 Viscoelastic Dampers

using basic trigonometry: cosωt 5 6

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 sin2 ωt

(A.17)

Substituting Equation (A.17) in Equation (A.16) and rearranging the terms, the following expression can be obtained:   sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   Fd;ve ðtÞ kve ud ðtÞ ud ðtÞ 2 1 12 5 cve ω u0 u0 cve ω u0

(A.18)

The above equation describes an inclined ellipse as shown in Fig. A.5. Similar to the viscous case, the maximum displacement and maximum force are out-ofphase, by an amount that can be adjusted according to viscoelastic properties (k and c from Equations (4.7) and (4.8), respectively). The energy dissipated by the VE damper in each cycle is the area under the force
displacement relationship, which also produces the same energy dissipated as the linear viscous element: ED 5 cve πωu20

(A.19)

As expected, the elastic portion does not contribute to the energy dissipation. Accordingly, the equivalent viscous damping ratio can be obtained, for an idealized SDOF system: ζ eq 5

cve 2ωm

(A.20)

Fd,ve (t) u 0cve ω 1 k ve 0.5 cve ω [(c /k ω)+1] 0.5+ {1 - 1/[(cve /k ve ω)+1]} ve ve k ve cve ω 1 k ve -1

1 1

ud u0

1 -1

[(cve /k ve ω)+1] 0.5

FIGURE A.5 Hysteretic behavior of a viscoelastic damper. Adapted from Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia.

977

978

APPENDIX A

where ̅ω is the oscillating circular frequency of the Kelvin element and m is the mass connected to its ends. The above equation can be rewritten in terms of elastic shear modulus, GE, and shear viscous damping constant, GC, as follows: ζ eq 5

cve ̅ω 2 cve ω GC ω 5 5 2kve 2GE 2ωkve

(A.21)

Looking at Equation (A.21), the term GE measures the energy stored and recovered for each cycle, and GC ω is the shear loss modulus that gives a measure of the energy dissipated during each cycle. Among the properties of viscoelastic material, it is common practice to define another measure of the energy dissipation capacity (i.e., the loss factor) (Mahmoodi, 1969). The loss factor, η, can be determined through experimental results and is related to the equivalent viscous damping constant defined previously. The loss factor can also be determined in terms of the shear storage and shear loss modulus, as follows: η5

GC ω GE

(A.22)

As defined by Equation (A.22), and when combined with Equation (A.21), the estimation of the damping constant can be shown as follows: cve 5

ηkve ω

(A.23)

A.3 HYSTERETIC DAMPER For the characterization of the damper performance, following variables are necessary (Fig. 4.31): yield force of the damper (Fy ), initial stiffness of the damper (k0 ), total force of the structure (Fmax ), and elastic stiffness of the structure (kE ). Considering the properties of the system idealized with the elasto-plastic curves without strain hardening (Fig. 4.31) the damper properties can be expressed ~ and the force as the stiffness ratio of damper (kd0) to structure stiffness (kS), k, ratio between the damper yield force (Fdy) and total structural force (FS), f~, as follows: kd0 k~ 5 kS

(A.24)

Fdy f~ 5 FS

(A.25)

Based on the hysteretic properties of the damper, the equivalent viscous damping can be estimated as defined in Equation (3.34).

A.3 Hysteretic Damper

The hysteretic energy dissipation can be expressed by:   ED 5 4Fdy ud;max 2 udy

(A.26)

where the structural displacement is considered equal to maximum displacement of damper (ud;max ), and udy is the yield displacement of damper. Instead, the strain energy is defined as: ES0 5

 1 kS ud;max 1 kd0 udy 2

(A.27)

The equivalent viscous damping can be found by substituting Equations (A.26) and (A.27) into Equation (3.60) as follows:   2Fy ud;max 2 udy  ζ eq 5  π kE ud;max 1 k0 udy

(A.28)

~ ~ S , and udy 5 ud;max f , the equivalent viscous dampSince Fdy 5 f~FS , kd0 5 kk k~ ing can be converted to:   f~ 2f~ 1 2 k~  ζ eq 5  ~ π 11f

(A.29)

The equivalent viscous damping can also be expressed based on bilinear hysteretic simplification (Fig. 4.32), as a function of postyield stiffness of the structure, kd0p 5 αd kd0 , and the displacement ductility, μ 5 ud;max =udy , as follows (Christopoulos and Filiatrault, 2006): ζ eq 5

2ð1 2 αd Þðμ 2 1Þ πμðαd μ 2 αd 1 1Þ

(A.30)

The general behavior of a hysteretic damper can be obtained by plotting the variation in equivalent viscous damping with k~ and f~ (Fig. A.6). From the figure, it is clear that as the damper stiffness increases, compared to the structure’s stiffness, the damping also increases. Moreover, an optimum value (the local maximum) can ~ This is demonstrated in Christopoulos and be determined for each stiffness ratio, k. Filiatrault (2006) for an SDOF system under harmonic excitation, where k~ 5 0:55 (Fig. A.7). In Fig. A.7, the response is shown based on three parameters: •

The amplitude of dynamic deformation u0 normalized by the equivalent static lateral displacement, ust , is given as: ust 5

•

F0 kS

(A.31)

where F0 is the amplitude of the equivalent applied load (Christopoulos and Filiatrault, 2006). The frequency ratio, σ, expressed as the ratio between the excitation circular frequency, ωe , and the structure natural frequency, ω, is as follows: ωe ω rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi kS 1 kd0 ω5 m σ5

(A.32) (A.33)

979

980

APPENDIX A

FIGURE A.6 Equivalent viscous damping vs. damper properties. Adapted from Kelly, T.E., 2001a. In-Structure Damping and Energy Dissipation
Design Guidelines. Holmes Consulting Group Ltd, New Zealand.

FIGURE A.7 SDOF response with hysteretic damper under harmonic excitation. Adapted from Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia.

•

The hysterically damped system’s parameter, Λhd , expressed as the ratio of the equivalent static lateral displacement and the displacement to activate the damper, udy can be given as follows: Λhd 5

ust kd0 F0 F0 5 5 k~ udy kS Fdy Fdy

(A.34)

A.4 Friction Damper

Based on Fig. A.7, the following conclusion has been drawn by Christopoulos and Filiatrault (2006): • •

• •

For Λhd . 0:85, the hysteretic damper has small influence on the structural frequency. For Λhd , 0:85, the hysteretic damper has a greater influence on the structural frequency and also provides a damping increase for the decreasing values of Λhd : For Λhd , 0:15, the system behaves as fully braced and this could lead to higher response amplitudes. For Λhd  0:45, the system exhibits a lowered response and it is therefore considered the optimum value.

opt For finding the optimum activation load, Fdy , a closed-form solution for an SDOF system subjected to harmonic excitation can be utilized (see Christopoulos and Filiatrault (2006) for a detailed discussion). Christopoulos and Filiatrault (2006) show how the optimum activation load is dependent on the frequency, the amplitude of the ground motion, and the structural properties.

A.4 FRICTION DAMPER The equivalent viscous damping for Coulomb friction can be simply modeled as shown in Fig. 4.63. The energy dissipated per cycle can be computed similar to viscous dampers in Equation (A.2). Consequently, the corresponding equivalent viscous damping constant cds;eq is as follows: cds;eq 5

4Fds πωud

(A.35)

Subsequently, the equivalent damping ratio can be computed as follows: cds;eq 2Fds ζ eq 5 pffiffiffiffiffiffi 5 πku 2 mk d

(A.36)

Typical friction cyclic behavior is shown in Fig. A.8 where the static force is higher than the kinetic force. This behavior, while being simply expressed by Equation (4.21), is based on complex material phenomena where the main factors influence the true contact area (see Kim et al. (2004)). For this reason, modeling of friction dampers is similar to that of an ideal elasto-plastic material (Fig. 4.31) in which the slip load can be considered equivalent to a yield force. This force
displacement relationship can be simply expressed by Equation (4.21) in which, after the slip load is overcome, the force remains constant until the reverse cycle. One of the major parameters to consider while designing friction dampers is related to the optimal slip load because it influences the response of the system, both in terms of period of vibration and energy dissipation of the friction damper

981

982

APPENDIX A

(Pall and Marsh, 1982). Moreover, dampers do not usually slip under wind loads. According to Pall and Tall (2004) the lower bound could be 130% of the wind shear and the upper bound could be 75% of the member yield shear force. Indeed, an optimal slip load gives the minimum response as shown in Fig. A.9. The variation in the slip load should not be more than 25% (Pall and Tall, 2004).

FIGURE A.8 Typical frictional force-sliding displacement relation. Adapted from Kim, H.-J., Christopoulos, C., Tremblay, R., 2004. Experiment characterization of boltstressed non-asbestos organic (NOA) material-to-steel interface. Report No. UT2004-3, Department of Civil Engineering, University of Toronto, Canada.

FIGURE A.9 Friction damper optimal slip load. Adapted from Pall, A.S., Pall, R.T., 2004. Performance-based design using pall friction dampers
an economical design solution. In: Proceedings of the 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada.

A.5 Tuned-Mass Damper

Another important aspect to take into consideration while dealing with this type of dampers is the possible variation in the coefficient of friction based on pressure, frequency, temperature, load dwell, corrosion of friction materials, contamination, surface wear, bolt preload, and travel length (see Christopoulos and Filiatrault (2006) for a detailed discussion). One of the major source of variation is frequency dependency. However, Muallu and Belev (2002) showed that a rotation friction damper equipped with a brass pad is almost frequency-independent. For this reason, the authors utilized the Coulomb law for friction modeling. This clearly allows the utilization of a bilinear force
displacement relationship given the behavior of the three components of the system, friction damper, primary frame, and brace.

A.5 TUNED-MASS DAMPER The basic theory behind TMD is discussed in this section. The description starts from the simplest case of two masses and two undamped springs under harmonic actions, and ends with the most advanced solution of multi-degrees-of-freedom tuned-mass dampers (MDOF TMD) coupled to multi-degrees-of-freedom damped structures under various actions. The section ends with a brief discussion about multiple TMD (MTMD).

A.5.1 TMDS SDOF THEORY The TMD equations of motion were first developed by Den-Hartog (1956) for an undamped structure and undamped TMD (Fig. A.10A) and for an undamped structure and damped TMD (Fig. A.10B), where the subscript d refers to the TMD properties. When damping is considered, in the structure, the fix point frequencies of the Den Hartog (1956) formulation are no longer valid. For this reason, different authors (Randall et al., 1981; Warburton, 1982; Warburton and Ayorinde, 1980; Tsai and Lin, 1993; and Abubakar and Farid, 2012) proposed a generalized formulation for both damped structure and TMD (Fig. A.11). However, given these complexities, solutions to the equations of motions are provided by numerical analysis or empirical charts. In the following text, the most generalized solutions are reviewed but interested readers should refer to Connor (2003) for a detailed explanation. Fig. A.11 shows the most general system with TMD mass, md , stiffness, kd , and viscous damping coefficient, cd . The general equations of motions (based on Section 3.1 (Chapter 3)), considering both (periodic) external force (F) and ground acceleration (u€ g ), can be expressed as follows: md u€ d 1 cd u_ d 1 kd ud 1 md u€ 5 2 md u€ g ðtÞ

(A.37)

mu€ 1 cu_ 1 ku 2 cd u_ d 2 kd ud 5 2 mu€ g ðtÞ 1 F ðtÞ

(A.38)

983

984

APPENDIX A

FIGURE A.10 TMD SDOF: (A) Undamped structure
undamped TMD and (B) undamped structure
damped TMD.

FIGURE A.11 TMD SDOF: Damped structure—damped TMD system.

The solution for a harmonic excitation (Equations (A.43) and (A.44)) can be given as (Connor, 2003): •

Undamped structure u5

u€ g0 m F0 Ops1 eiψ1 2 Ops2 eiψ2 k k

(A.39)

ud 5

u€ g0 m F0 Ops3 eiψ3 2 Ops4 eiψ4 k k

(A.40)

A.5 Tuned-Mass Damper

•

•

Damped structure u5

u€ g0 m F0 Ops6 eiψ6 Ops5 eiψ5 2 k k

(A.41)

ud 5

u€ g0 m F0 Ops8 eiψ8 Ops7 eiψ7 2 k k

(A.42)

Harmonic excitations u€ g 5 u€ g0 sinωe t

(A.43)

F 5 F0 sinωe t

(A.44)

where Ops are the amplification factors of the pseudo-static responses; ψ are the phase angles between the response and the excitation (see Connor (2003) for a detailed explanation); and ωe is the harmonic excitation vibration frequency. It is important to note that for a mass ratio (μ) less than 5% the amplification factors and phase shift for the external loading and ground motion are essentially equal. In literature, several expressions for determining the TMD optimal properties can be found (tuning and damping ratio) depending on whether the main structure is undamped or damped. In the case of undamped structure, the optimal parameters can be computed for different control response criteria and loading cases as shown in Table A.1 (see Constatinou et al. (1998) for a detailed discussion). In the table, f d;opt is the optimal natural frequency ratio and ζ d;opt is the optimal supplemental damping ratio. For more details about the different expressions, readers can refer to the references provided. In case the damping of the structure is also considered, the optimum tuning frequency of TMDs is strongly dependent upon the inherent damping, especially in base acceleration inputs. Instead, the optimum supplemental damping ratio is less sensitive to inherent damping (Tsai and Lin, 1993; Bakre and Jangid, 2007). A number of relations are available in literature, as shown in Table A.2, that have been mostly determined through curve fittings since the equations of motions cannot be solved in closed-form solution (for more details about the different expressions readers should refer to the references provided). As an alternative to the relations presented in Table A.2, design charts are also available. For example, Connor (2003) numerically developed curves for different levels of inherent damping of the main system, ζ i , and mass ratio, μ, for the optimum tuning frequency (Fig. A.12), the optimum damping ratio (Fig. A.13). To demonstrate the influence in the selection of different parameters, Connor (2003) has shown the response of a simple damped SDOF system with a 1% mass ratio TMD and 2% inherent damping ratio both under harmonic (Fig. A.14) and seismic excitations (Fig. A.15). While in the first case the response is greatly reduced, in the second one there could also be a negative effect, that is, response amplification. This behavior is due to the ineffectiveness of the TMD to impulsive loading and its capacity to reach resonance under random excitation.

985

Table A.1 Optimum Parameter of a TMD for the Undamped Structure Modeled as an SDOF System Under Various Excitations Excitation

Control Response Criteria

Optimum TMD Parameters

Applied to

Den Hartog (1956) and Rana and Soong (1998)

Harmonic

Main system

Ayorinde and Warburton (1980)

White-noise ground acceleration

Warburton (1982)

Harmonic

Main system

Displacement

1 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 2 μ=4   4ð1 1 μÞ 1 2 μ=2 rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3μ 8ð1 1 μÞ

Harmonic

Main system

Acceleration

rffiffiffiffiffiffiffiffiffiffiffiffi 1 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3μ   8 1 1 μ=2

Harmonic

Base

Displacement

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3μ   8ð1 1 μÞ 1 2 μ=2

Harmonic

Base

Acceleration

1 11μ

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3μ 8ð1 1 μÞ

Random acceleration

Main system

Random acceleration

Base

Rana and Soong (1998)

Harmonic

Base

Displacement

Chang and Qu (1998)

Harmonic

Main system

Equivalent damping ratio

Displacement

Frequency Ratio, f d;opt

Damping Ratio, ζ d;opt

Type Reference

1 11μ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 11μ

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 μ=2 11μ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 11μ 1 11μ

rffiffiffiffiffiffiffiffiffiffiffiffi 22μ 2

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 μ=2 11μ

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3μ 8ð1 1 μÞ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 1 3μ=4   4ð1 1 μÞ 1 1 μ=2 sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 2 μ=4   4ð1 1 μÞ 1 2 μ=2 vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi sffiffiffiffiffiffiffiffiffiffiffiffi u u 3μ 2 t 8ð1 1 μÞ 2 2 μ sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 1 3μ=4   4ð1 1 μÞ 1 1 μ=2

Connor (2003)

Harmonic

Main system

Displacement

Chritopulos and Filiatrault (2006)

Harmonic

Main system

Displacement

Bakre and Jangid (2007)

White-noise random

Main system

Displacement

White-noise random

Main system

Velocity

White-noise random

Main system

Force transmitted to base

White-noise random

Base

Displacement

White-noise randoma

Base

Displacement

White-noise randomb

Base

Displacement

White-noise randomc

Base

Displacement

Harmonic

Main system

Displacement

Harmonic

Base

Displacement

White-noise random

Base

Displacement

Hoang et al. (2008)

Leung and Zhang (2009)

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffi ffi  μ 3 2 μ=2   8ð1 1 μÞ 1 2 μ=2

1 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3μ 8ð11μÞ3

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 μ=2 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 1 3μ=4   4ð1 1 μÞ 1 1 μ=2 rffiffiffi μ 4

1 pffiffiffiffiffiffiffiffiffiffiffiffi 11μ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 μ=2 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 1 3μ=4   4ð1 1 μÞ 1 1 μ=2

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 2 μ=4   4ð1 1 μÞ 1 2 μ=2

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ð1 2 0:6μÞð1 1 μ2 Þ 11μ

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi μð1 1 2:5μ 1 2μ2 Þ 2ð1 1 2:7μÞ

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 λμ=6 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 2 μ=4   4ð1 1 μÞ 1 2 μ=2

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 2 μ=4   4ð1 1 μÞ 1 2 μ=2

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 μ=2 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 1 3μ=4   4ð1 1 μÞ 1 1 μ=2

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 11μ

3μ   8ð1 1 μÞ 1 2 μ=2

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 2 μ=4   4ð1 1 μÞ 1 2 μ=2 (Continued)

Table A.1 Optimum Parameter of a TMD for the Undamped Structure Modeled as an SDOF System Under Various Excitations Continued Excitation Type

Applied to

Harmonic

Main system

Displacement

Harmonic

Main system

Velocity

Harmonic

Main system

Acceleration

White-noise random

Base

Displacement

Reference Tuan and Shang (2014)

Marian and Giaralis (2015)

Control Response Criteria

Optimum TMD Parameters Frequency Ratio, f d;opt 1 11μ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 μ=2 11μ

b

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3μ 8ð11μÞ3 sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  ffi μ 3 1 3μ 1 0:625μ2

rffiffiffiffiffiffiffiffiffiffiffiffi 1 11μ

4ð2 1 μÞð11μÞ3 rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3μ 4ð2 1 μÞð1 1 μÞ

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   ffi μ 1 2 μ=4   4ð1 1 μÞ 1 2 μ=2

If ground frequency ratio λ 5 ωg =ω (ratio between ground frequency and system mass frequency) is equal to 1, that is, λ 5 1. If 1 , λ , 3. c If λ $ 3. a

Damping Ratio, ζ d;opt

Table A.2 Optimum Parameter of a TMD for the Damped Structure Modeled as an SDOF System Under Various Excitations Excitation Reference

Type

Applied to

Tsai and Lin (1993)

Harmonic

Base

Bakre and Jangid (2007)

White-noise random

Main system

White-noise random

Main system

White-noise random

Main system

White-noise random

Base

White-noise randoma

Base

White-noise randomb

Base

White-noise randomc

Base

Hoang et al. (2008)

Optimum TMD Parameters Frequency Ratio, ̅f d;opt ! pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   pffiffiffi 1 2 μ=2 pffiffiffi 1 1 2 2ζ 2i 2 1 2 2:375 2 1:304 μ 2 0:426μ ζ i μ 11μ   pffiffiffi pffiffiffi 2 3:73 2 16:903 μ 1 20:696μ ζ i 2 μ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  pffiffiffi 1 1 μ=2  pffiffiffi 1 2 0:5072 1 0:11169 μ 1 0:52223μ ζ i μ 1 11μ   pffiffiffi pffiffiffi 2 0:01518 1 0:31876 μ 1 0:23187μ ζ i 2 μ   pffiffiffi 1 pffiffiffi pffiffiffiffiffiffiffiffiffiffiffiffi 1 1:07059 2 5:6296 μ 1 12:1299μ ζ i μ 1 11μ   pffiffiffi pffiffiffi 10:4101 2 20:4405 μ 2 12:3364μ ζ i 2 μ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  pffiffiffi 1 1 μ=2  pffiffiffi 1 2 0:53101 1 0:07162 μ 1 0:73233μ ζ i μ 11μ   pffiffiffi pffiffiffi 1 1:22473 2 1:64174 μ 1 1:88277μ ζ i 2 μ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  pffiffiffi 1 2 μ=2  pffiffiffi 1 2 3:79441 1 9:87259 μ 2 15:2978μ ζ i μ 11μ   pffiffiffi pffiffiffi 1 2 13:6731 1 19:1284 μ 1 21:7049μ ζ i 2 μ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ð1 2 0:6μÞð1 1 μ2 Þ 2 0:7ζ i 11μ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 λμ=6 0:7ζ i 2 1 2 μ=2 11μ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 0:7ζ i 2 1 2 μ=2 11μ

Damping Ratio, ζ d;opt sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   3μ   1 0:151ζ i 2 0:17ζ i 2 8ð1 1 μÞ 1 2 μ=2   1 0:163ζ i 1 4:98ζ i 2 μ sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   μ 1 1 3μ=4   4ð1 1 μÞ 1 1 μ=2 rffiffiffi  pffiffiffi μ  pffiffiffi 1 0:00492 1 0:70913 μ 2 0:05096 ζ i μ 4   pffiffiffi pffiffiffi 1 0:64502 1 1:29223 μ 2 1:69405 ζ i 2 μ sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   μ 1 1 3μ=4   4ð1 1 μÞ 1 1 μ=2 sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   μ 1 2 μ=4   4ð1 1 μÞ 1 2 μ=2 pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi μ ð1 1 2:5μ 1 2μ2 Þ 2ð1 1 2:7μÞ sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   μ 1 2 μ=4   1 0:25μζ i 4ð1 1 μÞ 1 2 μ=2 sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   μ 1 2 μ=4   1 0:25μζ i 4ð1 1 μÞ 1 2 μ=2 (Continued)

Table A.2 Optimum Parameter of a TMD for the Damped Structure Modeled as an SDOF System Under Various Excitations Continued Excitation Reference

Type

Applied to

Leung and Zhang (2009)

White-noise random

Main system

Harmonic

Base

White-noise random

Base

Optimum TMD Parameters Frequency Ratio, ̅f d;opt pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  pffiffiffi 1 1 μ=2  pffiffiffi 1 2 0:5047 1 0:0764 μ 1 06023μ ζ i μ 1 0:3737ζ i 2 μ 11μ ! pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  1 2 μ=2 pffiffiffi pffiffiffi 2 1 1 2 2ζ i 2 1 1 2 2:3662 1 1:1252 μ ζ i μ 11μ   pffiffiffi pffiffiffi 1 2 4:8287 1 25 μ 1 35μ ζ i 2 μ pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  pffiffiffi 1 2 μ=2  pffiffiffi 1 2 4:9453 1 20:2319 μ 2 37:9419μ ζ i μ 11μ  pffiffiffi pffiffiffi 1 2 4:8287 1 25 μ ζ i 2 μ

If ground frequency ratio λ 5 ωg =ω (ratio between ground frequency and system mass frequency) is equal to 1, that is, λ 5 1. If 1 , λ , 3. c If λ $ 3. a

b

Damping Ratio, ζ d;opt sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   μ 1 1 3μ=4   4ð1 1 μÞ 1 1 μ=2 rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   1 0:1277ζ i 1 3μ8ð1 1 μÞ 1 2 μ=2   0:5277ζ i 1 2:3657ζ i 2 μ ffi sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   μ 1 2 μ=4   2 5:3024ζ i 2 μ 4ð1 1 μÞ 1 2 μ=2

A.5 Tuned-Mass Damper

FIGURE A.12 Optimum tuning frequency for TMD. Adapted from Connor (2003).

FIGURE A.13 Optimum damping ratio for TMD. Adapted from Connor (2003).

Despite using complicated numerical searching techniques for finding the optimal solution, as found in the literature explained above, the following simplified procedures review different types of external excitations—harmonic motion and random excitation.

991

992

APPENDIX A

FIGURE A.14 SDOF response to harmonic excitation (Connor, 2003).

FIGURE A.15 SDOF response to seismic excitation, El Centro (Connor, 2003).

A.5 Tuned-Mass Damper

A.5.2 TMDS SDOF SIMPLIFIED PROCEDURE FOR HARMONIC EXCITATION Connor (2003) proposed a simplified procedure for the case of damped structure and damped TMD (Fig. A.11) for an harmonic excitation force. Considering the governing equations of motions (Equations (A.37) and (A.38)) and a harmonic excitation (Equations (A.43) and (A.44)), the following applies: •

Primary mass ð1 1 μÞu€ 1 2ζ i ωu_ 1 ω2 u 5

•

F 2 μu€ d m

(A.45)

Tuned mass u€ 1 2ζ d ωd u_ d 1 ω2d ud 5 2 u€

(A.46)

The TMD properties (i.e., mass, stiffness, and damping) need to be selected in order to give the optimal solution based on the bare structure properties. Given a frequency of excitation, ωe (Equation (A.44)), the response of the system can be expressed as follows:   u 5 u0 sin ωe t 1 ψ1   ud 5 ud0 sin ωe t 1 ψ1 1 ψ2

(A.47) (A.48)

where u0 and ud0 are the system and damper displacement amplitudes, respectively; and ψ1 and ψ2 are the phase angles determined as shown in Equations (A.53) and (A.54), respectively. In order to understand the influence of the different parameters, the damper frequency is tuned to the building fundamental frequency: ωd 5 ω

(A.49)

Consequently, the damper stiffness can be related to the structure stiffness as follows: kd 5 μk

(A.50)

In this case the solution can be derived as: vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi F0 u 1 u u0 5  2 t kμ 2ζ i 1 1 μ 1 2ζ1

(A.51)

d

1 u0 2ζ d   2ζ i 1 1 tanψ1 5 2 2ζ d μ ud0 5

tanψ2 5 2

π 2

(A.52) (A.53) (A.54)

993

994

APPENDIX A

The above relationships show that the mass damper is 90 out of phase with the primary mass. In this case, there is no damper and the response can be simplified as follows: u0 5

  F0 1 k 2ζ i

δ1 5 2

π 2

(A.55) (A.56)

In terms of total damping, ζ T , the response can be expressed as:   F0 1 k 2ζ T sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi   μ 2ζ i 1 2 11 1 ζT 5 2 2ζ d μ u5

(A.57)

(A.58)

The total damping shows the influence of the mass and damper damping. When increasing the mass the total damping increases, while when increasing the damper damping, the total damping decreases. However, there is a limit on the mass that can be added and reducing the damping increases the relative motion of the damper as well (Equation (A.52)). In order to understand the influence of the different parameters, an example can be shown (Connor, 2003). Assuming a total damping (ζ T ) of 10% with an inherent damping (ζ i ) of 0% and combining Equation (A.58) with Equation (A.52) the following can be obtained: μ 2

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  2ffi ud0 11 5 0:1 u0

(A.59)

Assuming the response of the damper is an order of magnitude greater than the structure (i.e., ud0 5 10u0 ) (Connor, 2003), Equation (A.59) can be simplified as:   μ ud0  0:1 2 u0   1 μ  2ζ T ud0 =u0

(A.60) (A.61)

Substituting the value of ζ T 5 0:1 and ud0 =u0 5 10 the mass ratio becomes: μ

2ð0:1Þ 5 0:02 10

(A.62)

The additional mass only needs to be 2% in order to get 10% additional damping. However, the downside is the accommodation of the relative displacement of a damper mass that is 10 times larger than that of the primary mass.

A.5 Tuned-Mass Damper

Subsequently, the damping of the system from Equation (A.52) can be shown as follows: ζd 5

  1 u0 5 0:05 2 ud0

(A.63)

and the damper stiffness from Equation (A.50) is: kd 5 μk 5 0:02k

(A.64)

It is important to note that the spring stiffness connecting the mass of the damper to the main structure could be the stiffness of the pendulum action, in case the TMD is hung from a cable attached to the structure (see Section 4.1.2.1.1 (Chapter 4)).

A.5.3 TMDS SDOF SIMPLIFIED PROCEDURE FOR RANDOM EXCITATION Once optimal parameters of the TMD have been calculated in order to lower the dynamic amplification factor under harmonic loading (see Chapter 4), it is possible to predict the performance of the TMD under any dynamic loading, using a modal-spectral approach on 2D model, for presizing, or time-history approach on multi-SDOF system for a much accurate sizing. As a presizing, one can consider the structure as an SDOF with specific parameters based on the main mode contribution without TMD, coupled with the TMD (Fig. A.11). Coupling Equations (A.37) and (A.38), the overall system can be considered as a two-DOF system with an equation of motion expressed in Equation (3.28), as for an MDOF system. In this case, the required matrices and vectors are as follows:

 m 0 M5 0 md 

c 1 cd 2cd C5 cd 

2cd k 1 kd 2kd K5 2k d kd uðtÞ U ðtÞ 5 ud ðtÞ F 2 mu€ g F ðt Þ 5 ½  2md u€ g

(A.65)

The displacement vector U ðtÞ can be written in a modal basis [φ1 ; φ2 T , as expressed in Equation (3.37). Substituting the modal-based equations into Equation (3.38) and premultiplying that by t φm leads to the following modalbased equation:

q€m ðtÞ φm Mφm 1 q_m ðtÞ φm Cφm 1 qm ðtÞ φm Kφm 5 φm t

where m 5 1; 2.

t

t

t

FðtÞ 2 mu€g ðtÞ 2md u€g ðtÞ

 (A.66)

995

996

APPENDIX A

Equation (A.66) can be simplified as following applies from the modal theory (Chopra, 2007): φ1 Mφ2 5 0 φ1 Cφ2 5 0 t φ1 Kφ2 5 0

t

t

(A.67)

Substituting qm ðtÞ 5 eiωm t into Equation (A.66), modal vectors

T and their pulsations ωm , with m 5 1; 2, can be determined φm 5 φ1m φ2m without any load acting, in free vibration, as follows:   ω2m mφ21m 1md φ22m 2ðk1kd Þφ21m 1 2kd φ1m φ2m 2 kd φ22m 5 0

(A.68)

Herein, the elements of φm vectors can be chosen so that TMD displacement at the basis is equal to 1; that is, φ11 5 1, φ12 5 1. Therefore, Equation (A.68) can be simplified as follows:   ω2m m1md φ22m 2ðk1kd Þ 1 2kd φ2m 2 kd φ22m 5 0

(A.69)

with the pulsation ωm corresponding to m-th mode verifying the equations:   det 2 ω2m M 1 K  5 0

(A.70)

mmd ω4m 2 ½mkd 1 md ðkd 1 kÞω2m 1 kd k 5 0

(A.71)

Subsequently, the modal vectors and corresponding pulsations can be determined as follows: 2 kd φ1m ðm 5 1; 2Þ md ω2m 2 kd qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ðmkd 1 md ðkd 1 kÞÞ 7 ðmkd 1md ðkd 1kÞÞ2 2 4kd kmmd φ2m 5

ω21;2 5

2mmd

(A.72)

(A.73)

Substituting φ1 and φ2 with elements φ11 5 1, φ12 5 1 and those given by Equation (A.72) into Equations (3.48) and (3.52), the modal mass, stiffness, and damping associated to the m-th mode can be computed as:

 m 0 φm 5 m 1 md φ22m M m 5 t φm

0 md  k 1 kd 2kd φm 5 k 1 kd ð12φ2m Þ2 K m 5 t φm 2k k d d

 c 1 cd 2cd φm 5 c 1 cd ð12φ2m Þ2 C m 5 t φm 2cd cd

(A.74)

and the relative equations of motion in modal form can be rewritten as: M m q€ m ðtÞ 1 C m q_ m ðtÞ 1 K m qm ðtÞ 5 FðtÞ 2 mu€ g ðtÞ 2 φ2m md u€ g ðtÞ

(A.75)

Dividing the above equation by modal mass M m , the modal form of differential equations read q€ m ðtÞ 1 2ζ Tm ωm q_ m ðtÞ 1 ωm 2 qm ðtÞ 5

FðtÞ 2 mu€ g ðtÞ 2 φ2m md u€ g ðtÞ Mm

(A.76)

A.5 Tuned-Mass Damper

where total modal damping ratio for m-th mode can be obtained by substituting M m and C m from Equation (A.74) into Equation (3.53): ζ Tm 5

c 1 cd ð12φ2m Þ2  2 2ωm m1md φ22m

(A.77)

The modal-based solution of Equation (A.75) can be derived using the Duhamel’s integral according to Equation (3.22) for a viscously damped SDOF system with time-varying mechanical excitation as (Chopra, 2007): qm ðtÞ 5

1 M m ωDm

ðt

Fm ðτÞe2ζ Tm ωm ðt2τ Þ sin½ωDm ðt 2 τ Þdτ

(A.78)

0

where Fm ðτ Þ 5 FðτÞ 2 ðm 1 φ2m md Þu€ g ðτÞ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ωDm 5 ωm 1 2 ζ Tm 2

(A.79) (A.80)

The solution of Equation (A.78) can be expressed in terms of spectral response as a function of the total damping ratio (ζ Tm ) and the frequency (ωm ) both assigned to the m-th mode:     SAm;max ωm ; ζ Tm 5 ω2m SDm ωm ; ζ Tm φm     SDm;max ωm ; ζ Tm 5 jjuujdmax jmax 5 SDm ωm ; ζ Tm φm

(A.81) (A.82)

where SAm;max is the m-th mode maximum pseudo-acceleration; SDm;max is the maximum displacement; and SDm is the m-th mode spectral displacement given as:     SDm ωm ; ζ Tm 5 max qm ðtÞ 0#t#N

(A.83)

where m 5 1; 2 with the modal responses expressed in Equation (A.78). The modal values can then be combined to get the total displacement and acceleration through, for example, the complete quadratic combination (CQC): juj

jaj

cqc SD;cqc 5 judcqc jcqc ; SA;cqc 5 jad jcqc

(A.84)

These spectral curves allow to make a presizing of the TMD, since the TMD does not modify the frequency of the main modes much and contributes only to increase the modal damping. This is done defining a TMD performance factor, H, as follows: H5

  SDm;max ωm ; ζ Tm   SDm;max ωm ; ζ d

or

H5

  SAm;max ωm ; ζ Tm   SAm;max ωm ; ζ d

(A.85)

997

998

APPENDIX A

Knowing the performance factor, for example, in terms of acceleration, the designer can determine the damping modal ratio of the mode coupled with the TMD, ζ Tm , and finally the other TMD characteristics with Equation (A.77). An example can better explain how this methodology can be utilized. Considering the temporal curve of global wind forces (Fig. A.16) the wind spectrum can be computed for different total modal damping values (0.5, 2, 4, 7, 10, and 30%) as shown in Fig. A.17. Considering a structure with main mode at 0.26 Hz, a TMD with a mass ratio of 0.5% and with a supplemental damping of 10% in the TMD,

FIGURE A.16 Temporal curve of wind during time (wind tunnel model) in accordance with EC1 (CEN, 2010).

FIGURE A.17 Horizontal pseudo-acceleration spectrum generated from temporal wind forces acting on a tower (wind tunnel model).

A.5 Tuned-Mass Damper

tuned at 0.134 Hz, the CQC response of the two modes gives a decrease in acceleration of 45% compared to the solution without TMD (Fig. A.18).

A.5.4 TMDS MDOFS THEORY The most general description of the equations of motions for a structure equipped with TMD is based on MDOF systems with N degrees of freedom composed of spring, mass, and damping with only one component of displacement active

FIGURE A.18 Pseudo-acceleration on structure with or without TMD.

FIGURE A.19 TMD MDOFs system.

999

1000 APPENDIX A

(Fig. A.19). If the TMD is attached to N-th node (i.e., mass associated to top floor of building structure), the equations of motions are: mN u€ N 1 cN ðu_ N 2 u_ N21 Þ 1 kN ðu_ N 2 u_ N21 Þ 2 cd u_ d 2 kd ud 5 2 mN u€ g 1 FN md ðu€ d 1 u€ N Þ 1 cd u_ d 1 kd ud 5 2 md u€ g

(A.86) (A.87)

while for other nodes (i-th mass with i , N), the force equilibrium leads to mi u€ i 1 ci ðu_ i 2 u_ i21 Þ 1 ki ðui 2 ui21 Þ 2 ci11 ðu_ i11 2 u_ i Þ 2 ki11 ðui11 2 ui Þ 5 2 mi u€ g 1 Fi (A.88)

From Equation (A.86), one can consider the term ðcd u_ d 1 kd ud Þ of TMD as an external force applied to node N. For the sake of simplicity, Equations (A.86) and (A.88) can be expressed in a matrix format (based on Equation (3.28)) where the structural matrices are defined in Equations (3.29)
(3.31) and the force vector reads: 2

2 3 3 0 F1 6 F2 7 6 7 0 6 6 7 7 6 ^ 7 6 7 ^ 6 7 2 M1u€ g ðtÞ 1 6 7 6 Fi 7 6 7 0 6 6 7 7 4 ^ 5 4 5 ^ FN cd u_ d 1 kd ud

(A.89)

where 1 is called effective vector of order N with elements equal to unity. Accordingly, the modal form of equation of motion corresponding to m-th model is as follows: M m q€m ðtÞ 1 Cm q_m ðtÞ 1 K m qm ðtÞ 5 t φm F ðtÞ 2 t φm M1u€g ðtÞ 2 φNm ðcd u_d 1 kd ud Þ

(A.90)

where M m , K m , and C m are defined in Equations (3.46) and (3.52). To achieve the TMD effectiveness for a narrow frequency range, it is necessary to decide which modal resonant response is desired to be controlled. Assuming to control the first modal response, the relevant equation of motion can be obtained by substituting m 5 1 into Equation (A.90). Moreover, when the frequency of dynamic load is close to the first mode frequency ω1 , the overall response is dominantly affected by the first mode response. Hence, the displacement of node N (at which the TMD is attached) can be approximated by response contribution of the first mode, that is, uN  q1 φN1 . Thus, one can substitute q1 ðtÞ 5 uN ðtÞ=φN1 and its time derivatives into Equation (A.90), resulting in the dominant equations of motion of the MDOF system as follows: M~ 1 u€ N ðtÞ 1 C~ 1 u_ N ðtÞ 1 K~ 1 uN ðtÞ 5 F~ 1 ðtÞ 2 ðcd u_ d 1 kd ud Þ

(A.91)

M~ 1 5 M m =φ2N1

(A.92)

C~ 1 5 C m =φ2N1

(A.93)

where

A.5 Tuned-Mass Damper 1001

F~ 1 ðtÞ 5

t

K~ 1 5 K m =φ2N1

 φ1 FðtÞ 2 t φ1 M1u€g ðtÞ =φN1

(A.94) (A.95)

Equation (A.91) represents the equivalent equation of motion of the MDOF system identical to that of an equivalent SDOF system. Therefore, based on Equation A.95 and A.87 (for TMD), two equations of motion coupling the displacement of node N (where the TMD is attached) and of the TMD, ud , the dynamic response of the MDOF-TMD system can be simply derived similar to that of an SDOF-TMD system discussed is Section A.5.3. The spectral curves for the signal FðtÞ for various damping values can then be estimated for the two modes knowing the TMD parameters determined from the performance factor described in the previous section.

A.5.5 MTMDS THEORY The studies on multiple TMD started in the 1990s with the application to SDOF system by Xu and Igusa (1992), Igusa and Ku (1994), Kareem and Klime (1995), and Jangid (1995). The extension to MDOF systems was done by Chen and Wu (2001). Contrary to TMDs that are tuned to a single frequency, MTMDs are tuned to several building mode of vibrations. Therefore, the number of TMDs depends on the vibration modes to be suppressed (Kareem and Klime, 1995). The MTMDs mass ratio can be estimated assuming different masses for each TMD, as shown in the following equation (Bakre and Jangid, 2004): PNd μ5

j51

mdj

m

(A.96)

where, mdj is the mass of the j-th TMD. Fig. A.20 illustrates the main structure as an SDOF system with damping in which Nd TMDs are attached, resulting in a MTMD system. Subsequently, the design of MTMD can be carried out in a similar manner as previously shown for TMD. The designer can design a single unit or a group of TMDs (consisting if a few dampers) composing the MTMD in such a way that the damper frequency is tuned to the selected mode of vibration to suppress (Lewandowski and Grzymislawska, 2009). Similar to TMDs, the MTMDs’ optimal parameters (i.e., damping ratio, tuning frequency, and bandwidth frequency) can be estimated through expressions found in the literature. For example, Bakre and Jangid (2004) proposed explicit relationships, where the main structure is assumed as a damped SDOF system, while supported by Nd TMDs (see Fig. A.20). In this case, the tuning frequency of TMDs is expressed by: f d;opt 5

ωd;ave ω

(A.97)

1002 APPENDIX A

FIGURE A.20 Structural model of a main system with MTMD (Bakre and Jangid, 2004).

where ω is the frequency of the main system; and ωd;ave is the average frequency of the MTMD system composed by Nd TMD, as shown: PNd ωd;ave 5

j51

ωj

Nd

(A.98)

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi The natural frequency of each TMD (i.e., ωd;j 5 kd;j =md;j ) is given by:  

 Nd 1 1 ω ωd;j 5 ωd;ave 1 1 j 2 2 Nd 2 1

(A.99)

with the bandwidth frequency, ωd , equaling: ωd 5

ωd;Nd 2 ωd;1 ωd;ave

(A.100)

According to Equation (A.98), the natural frequencies of the MTMD system are linearly distributed over a specified bandwidth. Subsequently, the optimum parameters of the MTMD system, which are functions of the number of TMDs, Nd , are presented for three Nd values, in Table A.3 (Bakre and Jangid, 2004). The relationships shown in the table are considered valid for inherent damping in the range of 0 # ζ i # 0:1 and mass ratio , 0:1, which cover the practical applications of TMDs. The number of MTMDs range from 5 to 21.

Table A.3 Optimum Parameters of a MTMD System for the Damped Structure Modeled as an SDOF System Under Harmonic Base Acceleration (Bakre and Jangid, 2004). Number of TMDs (n) 21 11

5

Optimum TMD Parameters Frequency Ratio, ̅f d;opt   1:00008 f d;opt n511 pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 11μ  pffiffiffi pffiffiffi 1 0:69514 μ 2 0:00241 1 0:72152 μ 2 0:4397μ   pffiffiffi 1 2 2:6114 1 1:4065 μ 1 2:3114μ  pffiffiffi pffiffiffi pffiffiffi μζ I 1 2 2:6114 1 0:3886 μ 1 1:3947μ μζ I 2   0:996 f d;opt n511

Damping Ratio, ζ d;opt   0:909 ζ d;opt n511 sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 3μ pffiffiffi   1 0:79377 μ 8ð1 1 μÞ 1 2 μ=2   pffiffiffi 0:5474 1 0:1038 μ 2 0:4522μ  pffiffiffi pffiffiffi 1 0:4201 2 1:0844 μ 1 0:4666μ μζ I  pffiffiffi 2 pffiffiffi 1 0:6866 2 0:5744 μ 1 0:4966μ μζ I   1:38 ζ d;opt n511

Bandwidth Frequency, ̅ω d;opt   1:06821 ω d;opt n511   pffiffiffi pffiffiffi 3:19268 μ 0:42113 1 0:04479 μ 2 0:38909μ  pffiffiffi pffiffiffi 1 6:4616 2 26:6088 μ 1 33:1912μ μζ i 1  pffiffiffi pffiffiffi 2 10:7884 1 17:8412 μ 1 34:2912μ μζ i 2

  0:83 ω d;opt n511

1004 APPENDIX A

A.5.6 PENDULUM TMD Based on the equations of motion for a pendulum TMD, as shown in Equation (4.24), the following approximations are valid for small θ: ud 5 Lsinθ  Lθ

(A.101)

FTP  md

(A.102)

Consequently, the equation of motion becomes: md u€ d 1

md g ud 5 2 md u€ L

(A.103)

For the above equation the pendulum stiffness, kd;eq , and natural period, Td , can be determined as follows: md g L sffiffiffi L Td 5 2π g kd;eq 5

(A.104)

(A.105)

Since the period of the system depends on the length of the pendulum, it can be a limitation because the required length can be higher than the typical story height. This problem can be solved with a compound pendulum (Connor, 2003) in which the motion of the pendulum is magnified with an interior rigid link (Fig. A.21). The governing equation of motion for this system can be expressed as follows: md ðu€ 1 u€ 1 1 u€ d Þ 1

md g ud 5 0 L

(A.106)

FIGURE A.21 Schematic design of a compound pendulum. Adapted from Connor (2003).

A.6 Tuned Liquid Damper 1005

Since the rigid link moves in phase with the damper u1 5 ud , the equation of motion can be simplified as follows: md u€ d 1

md g md ud 5 2 u€ 2L 2

(A.107)

As shown, the equivalent stiffness is now mdg/2L and the device has a double effective length, compared to a simple pendulum case.

A.6 TUNED LIQUID DAMPER The equivalent mechanical model for TLD is shown in Section 4.1.2.2.2 (Chapter 4). The performance of a TLD can be expressed in charts that provide preliminary designs (Tait et al., 2008). Efficiency and robustness are used to define the performance as: • •

Efficiency, Ψ, referred to that of an equivalent TMD with the same mass as the water in the TLD Robustness, defined as the damper effectiveness to variations in the damper and structure parameters

The TLD effectiveness can be expressed as the amount of effective damping, ζ d;eff , that is added to the main structure. This is achieved when the response of a structure-TLD equals to that of an SDOF (Fig. A.22). Performance diagrams can then be drawn for both effective damping and relative motion between the main structure and the damper, ur , as shown in Fig. A.23 (Tait et al., 2008), as a function of the damping, mass, and tuning ratio.

FIGURE A.22 TLD modeling compared to an SDOF system.

1006 APPENDIX A

FIGURE A.23 Performance charts of equivalent linear TLD for 5 0:025 0:5 , f d , 1:5, and 0 , ζ d ð%Þ , 100: (a) ζ d;eff (%) and (b) ur . Adapted from Tait, M.J., 2008. Modelling and preliminary design of a structure-TLD system. Eng. Struct. 30, 2644
2655.

McNamara (1977) provides a closed-form solution for the effective damping for undamped primary structure as follows: ζ d;eff 5

f d μζ d  4 2 2 ð11μÞ2 f d 1 2ð1 1 μÞf d 2ζ 2d 2 1 1 f d μ 1 1

(A.108)

Subsequently, the total damping can be computed through this simplified expression (Luft, 1979): ζ T 5 0:8ζ i 1 ζ d;eff

(A.109)

In the case that an undamped structure is subjected to a white-noise force, the optimal values can be determined as follows (Warburton, 1982): ζ d;eff ;opt 5

1 4

sffiffiffiffiffiffiffiffiffiffiffiffiffiffi μ 1 μ2 1 1 3μ 4

11μ ur;opt 5 qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2μ 1 3μ2

(A.110)

(A.111)

The TLD efficiency can be expressed as follows (Tait et al., 2008): Ψ5

ζ d;eff 3 100 ζ d;eff ;opt

(A.112)

Instead, the robustness depends on the efficiency of tuning ratio, damper damping, and response amplitude. Since, ζ d and fd are, weakly and strongly,

A.6 Tuned Liquid Damper 1007

amplitude-dependent, respectively, the efficiency and robustness of a TLD vary with the amplitude. Tait et al. (2008) experimentally studied the efficiency and robustness of TLD for a wide range of parameters. The results show that TLD is highly efficient if properly tuned and fitted with suitable damping screens. A similar numerical model was utilized by Tait and Deng (2010) to study the influence of different tank geometries and their effect on the performance of TLDs. One of the most important parameters in the TLD tank design is the freeboard, G, which is the distance between the still liquid level and the top of the tank (Fig. A.24). The numerical studies have shown that a horizontal-cylindrical TLD is more robust and effective compared to other geometries. Moreover, Tait et al. (2005) state that if the TLD is utilized to reduce the motion in both directions, a simplified assumption of two independent 1D TLDs can be used. In case more complex geometries are needed, Love and Tait (2011) developed an equivalent linearized mechanical model for tank subjected to a 1D base excitation. In the case of a 2D response behavior this model has been updated by Love and Tait (2014). Alternately, to optimize the TLD performance, Samanta and Banerji (2010) proposed a modified TLD configuration in which the TLD is mounted on a platform attached to the top of the building with the help of a rigid rod and a flexible rotational spring (see Fig. A.25). The efficiency of such a TLD with an optimal rotational spring stiffness can be further increased, compared to the conventional TLD (i.e., rigidly connected to building), due to the rotational acceleration of the rod being in phase with the building top acceleration.

FIGURE A.24 TLD tank geometric study with screens. Adapted from Tait, M.J., Deng, X., 2010. The performance of structure-tuned liquid damper systems with different tank geometries. Struct. Control Health Monit. 17, 254
277.

1008 APPENDIX A

FIGURE A.25 Schematic view of modified-TLD-structure model (Samanta and Banerji, 2010).

A.7 TUNED LIQUID COLUMN DAMPER For tuned liquid column damper, Den Hartog (1956) developed an equivalent linear model which can be utilized in lieu of more complex nonlinear models. The optimal parameters’ expressions are derived from original work by Den Hartog (1956) similar to what was explained in previous sections. These expressions were updated by several authors for TLCDs, for example, Xu et al. (1992), Balendra et al. (1995), Abe et al. (1996), Won et al. (1996), Gao et al. (1997), Sadek et al. (1998), and Chang and Hsu (1999). An SDOF system equipped with a TLCD can be schematically represented as shown in Fig. A.26. The equations of motion can be written in this form (Sakai et al., 1989):

m 1 md Lmd

Lmd md





 

 

 u€ c 0 u_ k 0 F ðt Þ u 1 1 5 u€ d 0 cd u_ d 0 kd ud 0

(A.113)

where L 5 b=LTLCD is the length ratio with LLTCD 5 2dTL 1 b; md 5 ρF ATLCD LTLCD is the liquid mass of the TLCD with ATLCD as the cross-sectional area of the TLCD; cd;eq 5 2md ωd ζ d is the equivalent damping of the TLCD with pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ωd 5 2g=LTLCD as the damped frequency; kd 5 2ρF ATLCD g, the stiffness of the TLCD; and F ðtÞ is the external excitation. In the case of a harmonic excitation force (F ðtÞ 5 md Ae ω2e sinðωe tÞ), the nonlinear TLCD damping constant is estimated as follows (Yalla, 2001): cd 5

4ρF ATLCD ηd Ae ωe 3π

(A.114)

A.8 Base Isolation 1009

FIGURE A.26 Schematic representation of SDOF with TLCD. Adapted from Di Matteo, A., Lo Iacono, F., Navarra, G., Pirrotta, A., 2014. Optimal tuning of tuned liquid column damper systems in random vibration by means of an approximate formulation. Meccanica 50 (3), 795
808.

where ηd is the coefficient of head loss of the orifice, Ae is the amplitude of the harmonic motion, and ωe is the frequency of the harmonic motion. In a similar manner the damping can be derived for a random excitation as follows (Yalla, 2001): rffiffiffi 2 cd 5 ρ ATLCD ηd σv_ π F

(A.115)

where σv_ is the standard deviation of the liquid velocity. In order to solve the linear equivalent model, an iterative procedure is needed since it is dependent on the liquid velocity standard deviation, which is not an a priori assumption. Yalla (2001) has compared this equivalent linear approach and the nonlinear solution for various head-loss values. For the optimal parameters, similar to what was done for the TMD (Appendix A.6), a comparison between different forcing functions was done for an undamped structure by Yalla (2001) as shown in Table A.4.

A.8 BASE ISOLATION The basic principles of base isolation theory can be described with of an SDOF system (Fig. A.27). This can be extended to an MDOF base-isolated system, as an equivalent model of a building structure (Naeim and Kelly, 1999).

Table A.4 Comparison of Optimal Parameters TMD and TLCD External Force Random force on the structure

TMD, ̅f d;opt pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 μ=2 11μ

ζ d;opt sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  ffi 1 μ 1 1 3μ=4   2 ð1 1 μÞ 1 1 μ2

pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 μ=2 11μ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  ffi μ 1 2 μ=4 1   2 ð1 1 μÞ 1 2 μ2

TLCD, ̅f d;opt rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  ffi 2 1 1 μ 1 2 ̅L =2 11μ

ζ d;opt vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi u  u μ 1 1 μ 1 L 2 μ=4 ̅L u u  2  2u t ð1 1 μ Þ 1 1 μ 2 L μ 2

Random acceleration at the base

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi  2 1 1 μ 1 2 3L =2 11μ

vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ffi u  u μ 1 2 μ 1 3L 2 μ=4 u αu  2  2u t ð1 1 μÞ 1 1 μ 2 3L μ 2

A.8 Base Isolation 1011

FIGURE A.27 SDOF base-isolated system.

For an MDOF system with N degrees of freedom, the governing equation of motion subjected to a horizontal ground acceleration (u€ g ) in the horizontal direction can be expressed as (Naeim and Kelly, 1999):   M U€ ðtÞ 1 CU_ ðtÞ 1 KU ðtÞ 5 2 M1 u€ g ðtÞ 1 u€ b ðtÞ

(A.116)

where superstructure matrices (M, C, K) and vectors (U, 1) are defined in Equations (3.29) to (3.35); ub is the displacement of the base slab relative to the ground displacement (ug ); and 1 is the N-dimensional influence coefficient vector. Note that U ðtÞ is the displacement vector of order N, relative to the base slab. The matrix form of equations of motion for the building structure including the base isolation can be generalized as (Naeim and Kelly, 1999): 



      M U€ ðtÞ 1 C U_ ðtÞ 1 K U ðtÞ 5 2 M 1 u€ g ðtÞ

(A.117)

with 

M 5



m 1 mb M1





  c k 0 1T M ;C 5 b ;K 5 b 0 C 0 M



   1 u 1 5 ;U 5 b U 0

0 K

 (A.118) (A.119)

where mb , cb , and kb are the mass, damping, and stiffness of the base slab, respectively. Considering the postyielding behavior of base isolation, the governing equation of motion for the base slab can be expressed (neglecting the rotation and vertical deflection, based on Naeim and Kelly (1999)) as:       mb u€ g 1 u€ b 1 cb u_ b 1 αp0 k0 ub 1 1 2 αp0 k0 uh 1 1T M U€ 1 1u€ b 1 1u€ g 5 0

(A.120)

where αp0 is the postyield stiffness ratio; uh is a hysteretic component function of the time history; and k0 is the preyielding stiffness of the base slab. The hysteretic

1012 APPENDIX A

component is related to the base displacement with the following nonlinear relationship (Naeim and Kelly, 1999):   uh 5 A1 u_ b 2 A2 ju_ b juh juh jA4 21 1 A3 u_ b juh jA4 5 0

(A.121)

where A2 and A3 control the shape of the hysteretic loop; A1 controls the restoring force amplitude; and A4 controls the smoothness transition between elastic and plastic responses. These parameters relate to the yield displacement of the isolators as follows: udy 5

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi A1 A2 1 A3

(A.122)

Subsequently, the governing parameters for the dynamic behavior of the system can be derived as follows: sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi αp0 k0 ωb 5 T 1 M1 1 mb cb  ζ eq;b 5  T 2 1 M1 1 mb ωb

(A.123) (A.124)

where ωb is the frequency of the base slab; ζ eq;b is the equivalent viscous damping of the base slab; the period of the base slab Tb can be easily computed by Equation (3.10). To solve the complex governing equation of motion (given the fact that the seismic excitation is nonstationary in nature), several methods have been studied, including static correction, equivalent linearization (ELM), space state, and pseudo-excitation (PEM) (see Ma et al. (2014) for a comparison between each method). A detail description of these methods is outside the scope of this book.

References AASHTO, 2002. Guide Specifications for Seismic Isolation Design. American Association of State Highway and Transportation Officials, Washington, DC, 444 North Capitol Street, N.W. Suite 249. AASHTO, 2010. Guide Specifications for Seismic Isolation Design. American Association of State Highway and Transportation Officials, Washington, DC. Abbas, H., Kelly, J.M., 1993. A Methodology for Design of Viscoelastic Dampers in Earthquake-Resistant Structures. Earthquake Engineering Research Center, University of California, Berkeley. Abdullah, M.M., Richardson, A., Hanif, J., 2001. Placement of sensor/actuators on civil structures using genetic algorithms. Earthquake Eng. Struct. Dyn. 30 (8), 1167 1184. Abe, M., Fujino, Y., 1994. Dynamic characterization of multiple tuned mass dampers and some design formulas. Earthquake Eng. Struct. Dyn. 23 (8), 813 836. Abe, M., Kimura, S., Fujino, Y., 1996. Control laws for semi-active tuned liquid column damper with variable orifice opening. In: Proceedings of 2nd Int. Workshop on Structural Control, Hong Kong. Abubakar, I.M., Farid, B.J.M., 2012. Generalized Den Hartog tuned mass damper system for control of vibrations in structures. WIT Trans. State Art Sci. Eng. 59, 185 193. Achkire, Y., 1997. Active tendon control of cable-stayed bridges. Thesis. Universite´ Libre de Bruxelles, Bruxelles, Belgium. ACI, 2014. ACI-318: building code requirements for structural concrete and commentary. American Concrete Institute, Farmington Hills, MI. Adebe, D.Y., Choi, J., 2014. Analytical evaluation on hysteresis performance of circular shear panel damper. Int. J. Civil Environ. Struct. Construct. Arch. Eng. 8 (6), 744 750. Adebe, D.Y., Kim, J.W., Choi, J.H., 2013. Hysteresis characteristics of circular ripe steel damper using LY225. In: Steel Innovations Conference, Christchurch, New Zealand. Adeli, H., Saleh, A., 1997. Optimal control of adaptive/smart bridge structures. J. Struct. Eng. 123 (2), 218 226. Agrawal, A.K., Yang, J.N., Schmitendorf, W.E., Jabbari, F., 1997. Stability of actively controlled structures with actuator saturation. J. Struct. Eng. 126 (12), 1380 1387. Aguirre, M., Sa`nchez, A.R., 1992. Structural seismic damper. J. Struct. Eng. 118 (5), 1158 1171. AIJ, 1991. Guidelines for the Evaluation of Habitability to Building Vibration. Architectural Institute of Japan, Maruzen 82 (in Japanese). AIJ, 2000. Damping in Buildings. Architectural Institute of Japan, Maruzen 287. in Japanese. AIJ, 2004. Guidelines for the Evaluation of Habitability to Building Vibration. Architectural Institute of Japan, Maruzen 132. in Japanese. AIJ-GBV, 2004. Guidelines for the evaluation of habitability to building vibration, Architectural Institute of Japan, Maruzen, pp.132. (in Japanese). Aiken, I., 2006. Energy dissipation devices. 8NCEE, State of art, 100th Anniversary Earthquake Conference, April 18-22, San Francisco, California. Aiken, I.D., Kelly, J.M., 1990. Earthquake simulator testing and analytical studies of two energy-absorbing systems for multistory structures. Report No. UCB/EERC-90/03, EERC, University of California, Berkeley, CA.

1013

1014

References

Aiken, I.D., Kelly, J.M., Mahmoodi, P., 1990. The application of VE dampers to seismically resistant structures. In: Proceedings of 4th U. S. National Conference on Earthquake Engineering, Palm Springs, California, Vol. 3, pp. 459 468. Aiken, I.D., Douglas, K.N., Whittaker, A.S., Kelly, J.M., 1993. Testing of passive energy dissipation systems. Earthquake Spectra 9 (3), 335 370. AISC, 2016a. AISC-341: Seismic Provisions for Structural Steel Buildings. American Institute of Steel Construction, Chicago, Illinois. AISC, 2016b. AISC-360: Specification for Structural Steel Buildings. American Institute of Steel Construction, Chicago, Illinois. Akbay, Z., Aktan, H.M., 1990. Intelligent energy dissipation devices. In: Proceedings of the Fourth U.S. National Conference on Earthquake Engineering, Palm Springs, CA, 3, pp. 427-435. Alefeld, G., Herzberger, J., 1983. Introduction to Interval Computations. Academic Press, New York. Ali, M.M., Moon, K.S., 2007. Structural developments in tall buildings: current trends and future prospects. Arch. Sci. Rev. 50 (3), 205 223. Al-Kodmany, K., 2015. Tall buildings and elevators: a review of recent technological advances. Buildings 5, 1070 1104. Almufti, I., Willford, M., 2014. The REDit rating system: a framework to implement resilience-based earthquake design for new buildings. In: Tenth U.S. National Conference on Earthquake Engineering Frontiers of Earthquake Engineering, July 21 25, Anchorage, Alaska. Almufti, I., Krolicki, J., Chrowter, A., 2016. The resilience-based design of the 181 Fremont tower. Structure Magazine 42 46. June 2016. Aly, A.M., 2012. Proposed robust tuned mass damper for response mitigation in buildings exposed to multidirectional wind. Struct. Design Tall Special Buildings 23 (9), 664 691. Amini, F., Tavassoli, M.R., 2005. Optimal structural active control force, number and displacement of controllers. J. Struct. Eng. 27, 1306-1216. Andrew, J.P., Kaczmarczyk, S., 2011. Systems Engineering of Elevators. Elevator World, Inc, Mobile, Alabama. Anh, N.D., Nguyen, N.X., 2012. Extension of equivalent linearization method to design of TMD for linear damped systems. Struct. Control Health Monit. 19 (6), 565 573. Ankireddi, S., Yang, H.T.Y., 1996. Simple ATMD control methodology for tall buildings subject to wind loads. J. Struct. Eng. 22 (1), 83 91. ANSI/AISC 341-10, 2010. Seismic Provisions for Structural Steel Building. American Institute of Steel Construction, Chicago, Illinois. ANSYS, 2002. Software Manual. Ansys Inc, Southpointe, Canonsburg, PA. Aquino, R.E.R., 2006. Consideration of dynamic wind effects using a modified Davenport gust response factor formulation in the structural design and evaluation of selfsupporting antenna towers in the Philippines. Thesis (MS Civil Engineering), University of the Philippines, Diliman, Quezon City. Aquino, R.E.R., Tamura, Y., 2013a. On stick-slip phenomenon as primary mechanism behind structural damping in wind-resistant design applications. J. Wind Eng. Ind. Aerodyn. 115, 121 136. Aquino, R.E.R., Tamura, Y., 2013b. Structural damping estimation for wind-resistant design of 200m-high steel office building using the stick-slip model. J. Wind Eng. Ind. Aerodyn. 38 (135), 35 49.

References

Aquino, R.E.R., Tamura, Y., 2013c. Framework for structural damping predictor models based on stick-slip mechanism for use in wind-resistant design of buildings. J. Wind Eng. Ind. Aerodyn. 117, 25 37. Arakawa, T., Yamamoto, K., 2004 Frequencies and damping ratios of a high-rise building based on micro tremor measurements. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1 6. Arakawa, T., Sasaki, A., Satake, N., Suda, K., Tamura, Y., 2003. Damping evaluation using full-scale data of building in Japan. ASCE J. Struct. Eng. 129, 470 477. AIJ, 2011. Preliminary Reconnaissance Report of the 2011 Tohoku-Chico Taiheiyo-Oki Earthquake. Architectural Institute of Japan, Maruzen. Ariga, T., Kanno, Y., Takewaki, I., 2006. Resonant behavior of base-isolated high-rise buildings under long-period ground motions. Struct. Design Tall Spec. Build. 15 (3), 325 338. Aristizabal-Ochoa, J.D., 1986. Disposable knee bracing: improvement in seismic design of steel frames. J. Struct. Eng. ASCE 112 (7), 1544 1552. Arup, 2009. Sony City, Tokyo: a diagrid combined with base isolation. Arup J. 2, 49 51. AS, 1989. AS1170.2: Commentary of the Australian Wind Loading Code, Australia. AS, 2016. AS9100 Quality Management Systems Requirements for Aviation, Space and Defense Organization. Society of Automotive Engineers, Warrendale, PA. AS/NZS, 2001. AS/NZS 4671, Steel Reinforcing Materials, Australian and New Zealand Standard. AS/NZS, 2014. AS/NZS 1154, Structural Steel Welding, Australian and New Zealand Standard. AS/NZS, 2004. AS/NZS 1170.5 Supp 1:2004 structural design actions. Part 5 Earthquake actions. Standard New Zealand, Standard Australia. AS/NZS, 2011. AS/NZS 1170.2:2011 Structural design actions. Part 2 Wind actions. Standard New Zealand, Standard Australia. ASCE, 1999. ASCE Manual of Practice No. 67. American Society of Civil Engineering. ASCE, 2003. Seismic evaluation of existing buildings, American Society of Civil Engineers, ASCE/SEI 31-03, Reston, VA. ASCE, 2005. ASCE SEI7-05: Minimum Design Loads for Buildings and Other Structures. ASCE Publications, Reston, VA. ASCE, 2007. Seismic rehabilitation of existing buildings, American Society of Civil Engineers, ASCE/SEI 41-06, Reston, VA. ASCE, 2010. ASCE7-10: Minimum Design Loads for Buildings and Other Structures. American Society of Civil Engineers, Reston, VA. ASCE, 2013. ASCE41-13: Seismic Evaluation and Retrofit of Existing Buildings. American Society of Civil Engineers, Reston, VA. ASCE, 2015. Wind-Induced Motion of Tall Buildings: Design for Habitability. American Society of Civil Engineers, Reston, VA. ASCE, 2017a. ASCE7-16: Minimum Design Loads for Buildings and Other Structures. American Society of Civil Engineers, Reston, VA. ASCE, 2017b. ASCE41-17: Seismic Evaluation and Retrofit of Existing Buildings. American Society of Civil Engineers. Ashasi-Sorkhabi, A., Kristie, J., Mercan, O., 2014. Investigations of the use of multiple tuned liquid dampers in vibration control. Struct. Cong. 2004, 1185 1196.

1015

1016

References

Ashasi-Sorkhabi, A., Malekghasemi, H., Ghaemmaghami, A., Mercan, O., 2017. Experimental investigations of tuned liquid damper-structure interactions in resonance considering multiple parameters. J. Sound Vibrat. 388, 141 153. ATC, 1978. Tentative provisions for the development of seismic regulations for buildings, Applied Technology Council, ATC-3-06 report, Palo Alto, CA. ATC, 1983. Seismic retrofitting guidelines for highway bridges, Applied Technology Council, ATC-6-2 report, Palo Alto, CA. ATC, 1987. Evaluating the seismic resistance of existing buildings, Applied Technology Council, ATC-14 report, Redwood City, CA. ATC, 1996. ATC-40: Seismic evaluation and retrofit of concrete buildings. Redwood City, CA. ATC, 2010. PEER-ATC 72-1 Modeling and Acceptance Criteria for Seismic Design and Analysis of Tall Buildings. Applied Technology Council, Redwood City, California. ATC, 2016. ATC-114: Development of Accurate Models and Efficient Simulation Capabilities for Collapse Analysis to Support Implementation of Performance Based Seismic Engineering. National Institute of Standards and Technology. Athena, 2013. Eco-calculator for buildings. software and database overview. Athena Sustainable Materials Institute, Ontario, Canada. Ayorinde, E.O., Warburton, G.B., 1980. Minimizing structural vibrations with absorbers. Earthquake Eng. Struct. Dyn. 8, 219 236. Bagheri, S., Barghian, M., Saieri, F., Farzinfar, A., 2015. U-shaped metallic-yileding damper in building structures: seismic behaviour and comparison with a friction damper. Structures 3, 163 171. Baines, W.D., Peterson, E.G., 1951. An investigation of flow through screens. Trans. ASME 73, 467 480. Bairrao, R., Guerreiro, L., Barros, R.C., 2008. Shaking table tests on semi-active tuned mass and tuned liquid dampers. In: The 14 th World Conference on Earthquake Engineering October 12 17, Beijing, China. Bakre, S.V., Jangid, R.S., 2004. Optimum multiple tuned mass dampers for base-excited damped main system. Int. J. Struct. Stab. Dyn. 4 (4), 527 542. Bakre, S.V., Jangid, R.S., 2007. Optimum parameters of tuned mass damper for damped main system. Struct. Control Health Monit. 14 (3), 448 470. Balas, M.J., 1978. Feedback control of flexible structures. IEEE Trans. Automatic Control 673 679. AC-23. Balas, M.J., 1979. Direct velocity feedback control of large space structures. J. Guid. Control. Dyn. 2 (3), 252 253. Balas, M.J., 1982. Theory for distributed parameter systems. Control Dyn. Syst. 18, 361 421. Balazic, J., Guruswamy, G., Elliot, J., Pall, R.T., Pall, A., 2000. Seismic rehabilitation of justice headquarters building Ottawa, Canada. 12WCEE. Balendra, T., Lim, E.L., Lee, S.L., 1994. Ductile knee braced frames with shear yielding knee for seismic resistant structures. Eng. Struct. 16 (4), 263 269. Balendra, T., Sam, M.T., Liaw, C.Y., 1990. Diagonal brace with ductile knee anchor for aseismic steel frame. Earthquake Eng. Struct. Dyn. 19, 847 855. Balendra, T., Wang, C.M., Cheong, H.F., 1995. Effectiveness of tuned liquid column dampers for vibration control of towers. Eng. Struct. 17 (9), 668 675.

References

Banerji, P., Muradi, M., Shah, A.H., Popplewell, N., 2000. Tuned liquid dampers for controlling earthquake response of structures. Earthquake Eng. Struct. Dyn. 29 (5), 587 602. Bannantine, J.A., Comer, J.J., Handrock, J.L., 1990. Efficient large scale nonlinear transient analysis by finite elements. Int. J. Numer. Method Eng. 10, 579 596. Barber, T., Wen, Y., 1980. Seismic response of hysteretic degrading structures. Publ. Turk Natl. Comm. Earthquake Eng. 7, 457 464. Basu, B., Bursi, O.S., Casciati, F., Casciati, S., Del Grosso, A.E., Domaneschi, M., et al., 2014. A European Association for the Control of Structures joint perspective. Recent studies in civil structural control across Europe. Struct. Control Health Monit. 21 (12), 1414 1436. BBC, 1970. Building Code City of Boston. Building Department City Hall Boston, Massachusetts. BCJ, 2013. Introduction to Building Standard Law Japanese Building Codes and Building Control System, (written by Mr. Hasegawa Tomohiro), The Ministry of Land, Infrastructure, Transport and Tourism, Tokyo, Japan. Beck, J.L., Jennings, P.C., 1980. Structural identification using linear models and earthquake records. Earthquake Eng. Struct. Dyn. 8 (2), 145 160. Becker, T., Furukawa, S., 2010. Comparison of US and Japanese Codes and Practices for Seismically Isolated Buildings. 2010 Structure Congress, ASCE, Orlando, Florida. Becker, T.C., Yamamoto, S., Hamaguchi, H., Higashino, M., Nakashima, M., 2015. Application of isolation to high-rise buildings: a Japanese design case study through a US design code lens. Earthquake Spectra 31 (3), 151 1470. Bekdas, G., Nigdeli, S.M., 2013. Response of discussion on estimating optimum parameters of tuned mass dampers using harmony search. Eng. Struct. 54, 265 267. Benavent-Climent, A., 2010. A brace-type seismic damper based on yielding the walls of hollow structural sections. Eng. Struct. 32 (4), 1113 1122. Bentz, A., 2012. Dynamics of Tall Buildings: Full-Scale Quantification and Impacts on Occupant Comfort. Ph.D. thesis, Univ. of Notre Dame, Notre Dame, IN. Bentz, A., Kijewski-Correa, T., 2008. Predictive models for damping in buildings: the role of structural system characteristics. In: Structures Congress, 18th Analysis and Computation Specialty Conf., Structural Engineering Institute (SEI), Reston, VA. Bergman, L.A., Mc Farland, D.M., Hall, J.K., Johnson, E.A., Kareem, A., 1989. Optimal distribution of tuned mass dampers in wind sensitive structures. Proceedings of 5th ICOSSAR, New York. Bernal, D., Kojidi, S.M., Kwanm K., Do¨hler, M., 2012. Damping Identification in buildings from earthquake records. In: SMIP12 Seminar on Utilization of Strong-Motion Data Proceedings, Sacramento, California. Bernal, D., Do¨hler, M., Kojidi, S.M., Kwan, K., Liu, Y., 2015. First mode damping ratios for buildings. Earthquake Spectra 31 (1), 367 381. Bisch, P., Carvalho, E., Degee, H., Fajfar, P., Fardis, M., Franchin, P., et al., 2012. Eurocode 8: Seismic Design of Buildings Worked Examples. Publications Office of the European Union: Luxembourg. Blakeborough, A., Williams, M.S., Darby, A.P., Williams, D.M., 2001. The development of real-time substructure testing. R. Soc. Lond. 359, 1869 1891. Blakeley, R.W.G., 1979. Analysis and design bridges incorporating mechanical energy dissipating devices for earthquake resistance. In: Proceeding Workshop Earthquake Resistant Highway Bridges, ATC 6-1, California, USA, pp. 314 342.

1017

1018

References

Blandon, C.A., Priestley, M.J.N., 2005. Equivalent viscous damping equations for direct displacement based design. J. Earthquake Eng. 9 (2), 257 278. Blevins, R.D., 1984. Applied Fluid Dynamics Handbook. Macmillan, Agincourt, ON, Canada. Blume, J.A., 1970. The motion and damping of buildings relative to seismic response spectra. Bull. Seismol. Soc. Am. 60 (1), 231 259. BNBC, 1993. Bangladesh National Building Code. Housing and building Research Institute, Bangladesh and Testing Institute. BNBC, 2015. Bangladesh National Building Code. Housing and building Research Institute, Bangladesh and Testing Institute. Bondonet, G., Filiatraul, A., 1997. Frictional response of PTFE sliding bearings at high frequencies. J. Bridge Eng. ASCE 2 (4), 139 148. Bossens, F., Preumont, A., 2001. Active tendon control of cable-stayed bridges: a largescale demonstration. Earthquake Eng. Struct. Dyn. 30 (7), 961 979. Bouc, R., 1967. Forced vibration of mechanical systems with hysteresis. In: Proceedings of the Fourth Conference on Nonlinear Oscillation. Prague, Czechoslovakia, p. 315. Bouc, R., 1971. Mode`le mathe´matique d’hyste´re´sis: application aux syste`mes a` un degre´ de liberte´. Acustica (in French) 24, 16 25. Braz-Ce`sar, M., Carneiro De Barros, R., 2013. Passive control of civil engineering structures. In: Proceeding of the 4th International Conference on Integrity, Reliability and Failure, Funchal, Portugal. Brincker, R., Zhang, L.M., Anderson, P., 2000. Modal identification from ambient response using frequency domain decomposition. In: Proc. 18th International Modal Analysis Conference (IMAC), San Antonio, TX. Brown, A.S., Ankireddi, S., Yang, T.Y., 1999. Actuator and sensor placement for multiobjective control of structures. J. Struct. Eng. 125 (7), 757 765. Bruneau, M., Reinhorn, A., 2006. Overview on resilience concept. In: Proceedings of the 8th U.S. National Conference on Earthquake Engineering, San Francisco, California. BS 5950-1, 1990. British standard- structural use of steelwork in building. Code of practice in design for simple and continuous construction: hot-rolled sections. BS 8110-2, 1985. British standard- structural use of concrete. Code of practice for special circumstances. BS ISO, 2010. BS ISO 4866 : 2010. Mechanical vibration and shock—Vibration of fixed structures—Guidelines for the measurement of vibrations and evaluation of their effects on structures. The International Organization for Standardization. Buckle, I.B., Mayes, R.M., 1990. Seismic isolation: history, application, and performancea world view. Earthquake Spectra 6 (2), 161 201. Burton, M.D., 2006. Effects of Low-Frequency Wind Induced Building Motion on Occupant Comfort. PhD Thesis, The Hong Kong University of Science and Technology, Hong Kong. Burton, M.D., Kwok, K.C.S., Hitchcock, P.A., Denoon, R.O., 2006. Frequency dependence of human response to wind-induced building motion. J. Struct. Eng. ASCE 132 (2), 296 303. Burton, M.D., Kwok, K.C.S., Abdelrazaq, A., 2015. Wind-induced motion of tall buildings: designing for occupant comfort. Int. J. High-Rise Build. 4 (1), 1 8.

References

Caccese, V., Elgaaly, M., Chen, R., 1993. Experimental study of thin steel-plate shear walls under cyclic load. J. Struct. Eng. 119 (2), 573 587. Cammelli, S., Li, Y.F., Mijorski, S., 2016. Mitigation of wind-induced accelerations using Tuned Liquid Column Dampers: experimental and numerical studies. J. Wind Eng. Ind. Aerodyn. 155 (2016), 174 181. Cao, H., Li, Q.S., 2004. New control strategies for active tuned mass damper systems. Comput. Struct. 82, 2341 2350. Carr, A.J., RUAUMOKO, 2004. Computer Program Library, Department of Civil Engineering. University of Canterbury, Christchurch, New Zealand. Carr, A., 2007. Ruaumoko Theory Manual. University of Canterbury. Casciati, F., De Stefano, A., Matta, E., 2003. Simulating a conical tuned liquid damper. Simul. Model. Pract. Theory 11 (5 6), 353 370. Casciati, F., Rodellar, J., Yildirim, U., 2012. Active and semi-active control of structures theory and applications: a review of recent advances. J. Intell. Mater. Syst. Struct. 23 (11), 1181 1195. Caughey, T.K., O’Kelly, M.E.J., 1965. Classical normal modes in damped linear dynamic systems. J. Appl. Mech. (ASME) 32, 583 588. CBC, 2001. California Building Code. California Building Standards Commission. CBC, 2010. California Building Code, California Building Standards Commission. CCSM, 2004. Code for construction of shanghai municipality. Design of Steel Concrete Mixed Structure for High-Rise Building. Procedures, Shanghai. Celebi, M., 1996. Comparison of damping in buildings under low amplitude and strong motions. J. Wind. Eng. Ind. Aerodyn. 59 (2), 309 323. Celebi, M., Phan, L.T., Marshall, R.D., 1993. Dynamic characteristics of five tall buildings during strong and low-amplitude motions. Struct. Des. Tall Build. 2 (1), 1 15. CEN, 2003. Eurocode 8 Part 1: General rules, Seismic Actions and Rules for Buildings. European Committee for Standardization, Brussels. CEN, 2004. EN 1337 Structural Bearings: Sliding Elements. European Committee for Standardization, Brussels, Belgium. CEN, 2005. EN 10025 European Structural Steel Standard. European Committee for Standardization, Brussels, Belgium. CEN 1999, 2007. Eurocode 9: Design of aluminium structures. CEN. CEN, 2009. EN 15129 Anti-seismic Devices. European Committee for Standardization. CEN, 2010. Eurocode 1: Actions on Structures Part 1-4: General Actions Wind Actions. ENV 1991-1-4:2005. European Committee for Standardization. CEN, 2012. EN 1090:2012, Part 1: Requirements for Conformity Assessment of Structural Components. European Committee for Standardization, Brussels. CSA, 2004. CSA A23.3-04: Design of Concrete Structures. Standards Council of Canada. Cha, Y.-J., Agrawal, A.K., Kim, Y., Raich, A.M., 2012. Multi-objective genetic algorithms for cost-effective distributions of actuators and sensors in large structures. Exp. Syst. Appl. 39, 7822 7833. Chan, R.W.K., Albermani, F., 2008. Experimental study on steel slit damper for passive energy dissipation. Eng. Struct. 30, 1058 1066. Chang, C.C., 1999. Mass dampers and their optimal designs for building vibration control. Eng. Struct. 21, 454 463. Chang, C.C., Henry, T.Y., 1995. Control of buildings using active tuned mass dampers. J. Eng. Mech. 121, 355 366.

1019

1020

References

Chang, C.C., Hsu, C.T., 1999. Control performance of liquid column vibration absorber. Eng. Struct. 20 (7), 580 586. Chang, C.C., Qu, W.L., 1998. Unified dynamic absorber design formulas for wind-induced vibration control of tall buildings. Struct. Design Tall Build. 7 (2), 147 166. Chang, K.C., Soong, T.T., Oh, S.T., Lai, M.L., 1992. Effect of ambient temperature on viscoelastically damped structure. J. Struct. Eng. 118 (7), 1955 1973. Chang, K.C., Soong, T.T., Oh, S.T., Lai, M.L., 1991. Seismic response of a 2/5 scale steel structure with added viscoelastic damper. Technical Report NCEER-91-0012, National Center for Earthquake Engineering Research, Buffalo, NY. Chang, K.C., Soong, T.T., Oh, S.T., Lai, M.L., 1995. Seismic behavior of steel frame with added viscoelastic dampers. J. Struct. Eng. 121 (10), 1418 1426. Chang, S., Tsai, K., Chen, K., 1998. Improved time integration for seudodynamic tests. Earthquake Eng. Struct. Dyn. 27, 711 730. Chang, Y., 2015. Analytical and Experimental Investigations of Modified Tuned Liquid Dampers (MTLDs). PhD Thesis. University of Toronto. Charleson, A., 2008. Seismic Design for Architects Outwitting the Quake. Architectural Press, Elsevier, Oxford, UK. Charleson, A., Guisasola, A., 2017. Seismic Isolation for Architects. Taylor&Francis, Routledge. Charleson, A.W., Allaf, N.J., 2012. Costs of base-isolation and earthquake insurance in New Zealand. In: Proceedings of the 2012 New Zealand Society of Earthquake Engineering (NZSEE) Conference, Christchurch, 13 15 April 2012. Charney, F.A., 2006. Unintended consequence of modeling damping in structures: Rayleigh damping. In: Procedures of 17th Analysis and Computation Specialty Conference, American Society of Civil Engineers. Chase, J.G., Mulligan, K.J., Gue, A., Alnot, T., Rodgers, G., Mander, J.B., et al., 2006. Reshaping hysteretic behaviour using semi-active resettable device dampers. Eng. Struct. 28 (10), 1418 1429. Chen, G., Chen, C., 2004. Semiactive control of the 20-story benchmark building with piezoelectric friction dampers. J. Eng. Mech. 130 (4), 393 400. Chen, Y., Peng, C., Xue, H., Ma, L., 2008. The Function and Economic of Fluid Viscous Dampers to Reduce Seismic and Wind Vibration in High-rise Buildings. Blue lake International Paper. Chen, G., Wu, J., 2001. Optimal placement of multiple tune mass dampers for seismic structures. J. Struct. Eng. 127 (9), 1054 1062. Chen, P.W., Robertson, L.E., 1972. Human perception thresholds of horizontal motion. ASCE J. Struct. Div. 98 (ST8), 1681 1695. Chen, S.H., Wu, J., 2004. Interval optimization of dynamic response for structures with interval parameters. Comput. Struct. 82 (1), 1 11. Chen, S.H., Ma, L., Meng, G.W., Guo, R., 2009a. An efficient method for evaluating the natural frequency of structures with uncertain-but-bounded parameters. Comput. Struct. 87 (9 10), 582 590. Chen, Y.Q., Cao, T.Z., Ma, L.Z., 2009b. Structural vibration passive control and economic analysis of a high-rise building in Beijing. Earthquake Eng. Eng. Vib. 8, 561 568. Cheng, P., Yongqi, C., 2015. The Selection and a Practical Optimization for Fluid Viscous Damper in High-Rise Building. Cheng, F.Y., Jiang, H., Lou, K., 2008. Smart Structures: Innovative Systems for Seismic Response Control. CRC Press, Taylor & Francis Group, Boca Raton, FL.

References

Chey, M.H., 2000. Parametric Control of Structures using a Tuned Mass Damper System under Earthquake Excitations. ME Thesis, University of Canterbury, Christchurch, New Zealand. Chey, M.H., 2007. Passive and Semi-Active Tuned Mass Damper Building Systems. Ph.D. Thesis, University of Canterbury, Christchurch, New Zealand. Chey, M.-H., Chase, J.G., Mander, J.B., Carr, A.J., 2010. Semi-active tuned mass damper building systems: design. Earthquake Eng. Struct. Dyn. 39, 119 139. Chia-Ming, C., Spencer, B.F., 2010. Active base isolation of buildings subjected to seismic excitations. Earthquake Eng. Struct. Dyn. 39, 1493 1512. Cho, B.-H., Jo, J.-S., Joo, S.-J., Kim, H., 2012. Dynamic parameter identification of secondary mass dampers based on full-scale tests. Struct. Control Health Monit. 27 (3), 218 230. Choi, H., Kim, J., 2010. New installation for viscoelastic dampers using cables. Canad. J. Civil Eng. 37 (9), 1201 1211. Choi, H.S., Joseph, L., 2012. Outrigger system design considerations. Int. J. High-Rise Build. 1 (3), 237 246. Choi, I.R., Park, H.G., 2009. Steel plate shear walls with various infill plate designs. J. Struct. Eng. ASCE 135 (Issue 7), 785 796. Choi, Y.-T., Wereley, N.M., 2009. Self-powered magnetorheological dampers. J. Vib. Acoust. 131 (4), 044501-1:5. Chooi, W.W., Oyadiji, S.O., 2008. Design, modelling and testing of magnetorheological (MR) dampers using analytical flow solutions. Comp. Struct. 86 (3 5), 473 482. Chopra, A.K., 2012. Dynamics of Structures: Theory and Application to Earthquake Engineering. Prentice Hall, Upper Saddle River, New Jersey. Chopra, A.K., McKenna, F., 2016. Modeling viscous damping in nonlinear response history analysis of buildings for earthquake excitation. Earthquake Eng. Struct. Dyn. 45 (2), 193 211. Christenson, R., Lin, Y.Z., Emmons, A., Bass, B., 2008. Scale experimental verification of semiactive control through real-time hybrid simulation. Struct. Eng. 134 (4), 522 534. Christenson, R.E., Spencer Jr., B.F., Hori, N., Seto, K., 2003. Coupled building control using acceleration feedback. Comp. Aided Civil Infrastruct. Eng. 18 (1), 4 18. Christopoulos, C., 2002. Self-Centering Post-Tensioned Energy Dissipating (PTED) Steel Frames for Seismic Regions. Ph.D. Thesis, Department of Structural Engineering, University of California, San Diego, CA. Christopoulos, C., Filiatrault, A., 2006. Principles of Passive Supplemental Damping and Seismic Isolation, IUSS Press, Istituto Universitario di Studi Superiori di Pavia. Christopoulos, C., Montgomery, M., 2013. Viscoelastic coupling damper (VCD) for enhanced dynamic performance of high-rise buildings. Earthquake Eng. Struct. Dyn. 12 (15), 2217 2233. Christopoulos, C., Filiatral, A., Folz., B., Uang, C.-M., 2002a. Post-tensioned energy dissipating connections for moment-resisting steel frames. ASCE J. Struct. Eng. 128 (9), 1111 1120. Christopoulos, C., Filiatrault, A., Folz, B., 2002b. Seismic response of self-centering hysteretic SDOF systems. Earthquake Eng. Struct. Dyn. 31 (5), 1131 1150. Christopoulos, C., Filiatrault, A., Folz, B., Uang, C.M., 2002a. Post-tensioned energy dissipating connections for moment-resisting steel frames. ASCE J. Struct. Eng. 128 (9), 1111 1120.

1021

1022

References

Christopoulos, C., Filiatrault, A., Folz, B., 2002b. Seismic response of self-centering hysteretic SDOF systems. Earthquake Eng. Struct. Dyn. 31 (5), 1131 1150. Chung, L., Reinhorn, A., Soong, T., 1988. Experiments on active control of seismic structures. J. Eng. Mech. 114 (2), 241 256. Clark, A.J., 1988. Multiple passive tuned mass dampers for reducing earthquake induced building motion. In: Proc., 9th World Conf. on Earthquake Engineering, Vol. 5, International Association for Earthquake Engineering, Tokyo-Kyoto, Japan, pp. 779 784. Clark, P., Aiken, I., Kasai, K., Ko, E., Kimura, I., 1999a. Design procedures for buildings incorporating hysteretic damping devices. Proc. 68th Annual Convention, Structural Engineers Association of California, Santa Barbara, California, 355 371. Clark, P.W., Aiken, I.D., Nakashima, M., Miyazaki, M., Midorikawa, M., 1999b. The 1995 Kobe (Hyogo-ken Nanbu) earthquake as a trigger for implementing new seismic design technologies in Japan. Lessons Learned Over Time, Learning From Earthquakes, Volume III. Earthquake Engineering Research Institute, Oakland, California. Clark, P.W., Aiken, I.D., Nakashima, M., Miyazaki, M., Midorikawa, M., 1999c. New Design Technologies: Lesson Learned Over Time, Learning From Earthquakes. Volume III. Earthquake Engineering Research Institute, Oakland, California. Clough, R., Penzien, J., 2003. Dynamics of Structures. Computers & Structures, Berkeley, CA. CNR, 2008. CNR DT 207-2008: Istruzioni per la Valutazione delle Aziioni e degli Effetti del Vento sulle Costruzioni, Rome, Italy (in Italian). Cole, R.J., Kernan, P.C., 1996. Life-cycle energy use in office buildings. Build. Environ. 31 (4), 307 317. Comber, M.V., Poland, C., Sinclair, M., 2012. Environmental Impact Seismic Assessment: Application of Performance-Based Earthquake Engineering Methodologies to Optimize Environmental Performance. Proceedings of Structure Congress, Chicago, pp. 910 921. Comsol, 2016. Comsol Multiphysics Version 5.2. Burlington, MA, USA. Connor, J.J., 2003. Introduction to Structural Motion Control. MIT-Prentice Hall Series on Civil, Environmental, and System Engineering. Constantinou, M.C., Symans, M.D., 1993. Experimental study of seismic response of buildings with supplemental fluid dampers. Struct. Design Tall Build. 10 (2), 93 132. Constantinou, M.C., Tadjbakhsh, I.G., 1985. Optimum characteristics of isolated structures. J. Struct. Eng. ASCE 111 (12), 2733 2750. 1985. Constantinou, M.C., Mokha, A., Reinhorn, A.M., 1990. Teflon bearings in base isolation. II: modeling. J. Struct. Eng. 116 (2), 455 474. Constantinou, M.C., Symans, M.D., Tsopelas, P., Taylor, D.P., 1993. Fluid Viscous Dampers in Applications of Seismic Energy Dissipation and Seismic Isolation, ATC 17-1, ATC. Constantinou, M.C., Soong, T.T., Dargush, G.F., 1998. Passive energy dissipation systems for structural design and retrofit. Monograph No. 1, Multidisciplinary Center for Earthquake Engineering Research, University at Buffalo, NY. Constantinou, M.C., Whittaker, A.S., Kalpakidis, Y., Fenz, D.M., Warn, G.P., 2007. Performance of Seismic Isolation Hardware under Service and Seismic Loading. MCEER-07-0012, Multidisciplinary Center for Earthquake Engineering Research, Buffalo, N.Y.

References

Cousins, W.J., Porritt, T.E., 1993. Improvements to lead-extrusion damper technology. Bull. N.Z. Natl. Soc. Earthquake Eng. 26, 342 348. Cousins, W.J., Robinson, W.H., McVerry, G.H., 1991. Recent developments in devices for bae isolation. In: Proceeding Pacific Conference on Earthquake Engineering, New Zealand, 2, pp. 221 232. Crouse, C.B., Leyendecker, E.V., Somerville, P.G., Power, M., Silva, W.J., 2006. Development of seismic ground-motion criteria for the ASCE/SEI 7 standard. In: Proceedings of 8 th U.S. National Conference on Earthquake Engineering, April 18 22, San Francisco, California, on CD. Cruz, C., 2017. Evaluation of Damping Ratios Inferred from the Seismic Response of Buildings. PhD Thesis Stanford University, California. Cruz, C., Miranda, E., 2016. Evaluation of damping ratios for the seismic analysis of tall buildings. J. Struct. Eng. 143 (1), 04016144. Cruz, C., Miranda, E., 2017a. Evaluation of the Rayleigh damping model for buildings. Eng. Struct. 138, 324 336. Cruz, C., Miranda, E., 2017b. Evaluation of soil-structure interaction effects on the damping ratios of buildings subjected to earthquakes. Soil Dyn. Earthquake Eng. 100, 183 195. CSA, 2009. CSA S16-09: Design of Steel Structures. Standards Council of Canada. CSA, 2004. CSA A23.3-04: Design of Concrete Structures. Standards Council of Canada. CSA, 2014. CSA A23.3-14: Design of Concrete Structures. Standards Council of Canada. CSI, 2011. ETABS: Integrated Analysis, Design and Drafting of Building Systems. Computers and Structures Inc, Berkeley, CA. CSI, 2016a. ETABS: Integrated Analysis, Design and Drafting of Building Systems. Computers and Structures Inc, Berkeley, CA. CSI, 2016b. SAP2000: Integrated Solution for Structural Analysis and Design. Computers and Structures Inc, Berkeley, CA. CSI, 2016c. An overview of the tuned mass damper and guidelines for modelling the device. SAP2000. Berkeley, CA. CTBUH, 2008. Recommendations for the seismic design of high-rise buildings. In: Report of the Working Group on the Seismic Design of Tall Buildings, prepared by Willford, M., Whittaker, A.S., and Klemencic, R., edited by Wood, A., for the Council on Tall Buildings and Urban Habitat, Chicago, Illinois. CTBUH, 2010. Tall buildings in numbers. CTBUH J. (Issue II). D’Aniello, M., Gu¨neyisi, E.M., Landolfo, R., Gu¨neyisi, E., 2014. Seismic Response of Steel Frames with Knee Braces. In: Proceedings of EUROSTEEL 2014 the 7th European Conference on Steel and Composite Structures, on stick, Napoli, Italia. Daniel, Y., Lavan, O., 2015. Optimality criteria based seismic design of multiple tunedmass-dampers for the control of 3D irregular buildings. Earthquake Struct. 8 (1), 77 100. Daraphet, M., 2013. Tall Buildings: Imaginative Fac¸ades Solutions. CTBUH publication. Dargush, G.F., Soong, T.T., 1995. Behavior of metallic plate dampers in seismic passive energy dissipation systems. Earthquake Spectra 11 (4), 545 568. Datta, T.K., 2003. A state-of-the-art review on active control of structures. ISET J. Earthquake Technol. 40 (1), 1 17. Davenport, A.G., Hill-Carroll, P., 1986. Damping in tall buildings: its variability and treatment in design. In: Isyumov, N., Tschanz, T. (Eds.), Building Motion in Wind. ASCE, New York, pp. 42 57.

1023

1024

References

De Angelis, M., Perno, S., Reggio, A., 2012. Dynamic response and optimal design of structures with large mass ratio TMD. Earthquake Eng. Struct. Dyn. 41 (1), 41 60. Debbarma, R., Chakraborty, S., Ghosh, S., 2010. Unconditional reliability-based design of tuned liquid column dampers under stochastic earthquake load considering system parameters uncertainties. J. Earthquake Eng. 14 (7), 970 988. De la Llera, J., Esquerra, C., Almazan, J.L., 2004. Earthquake behavior of structures with Copper Energy Dissipaters. Earthquake Eng. Struct. Dyn. 33, 329 358. De Silva, C.W., 2007. Vibration Monitoring, Testing, and Instrumentation. Taylor & Francis. CRC Press, Boca Raton, FL. Den Hartog, J.P., 1956. Mechanical Vibrations. McGraw-Hill Book Company, New York. Deng, X., Tait, M.J., 2009. Theoretical modelling of TLD with different tank geometries using linear long wave theory. ASME J. Vib. Acoust 131 (4), 041014. Deng, K., Pan, P., Li, W., Xue, Y., 2015. Development of a buckling restrained shear panel damper. J. Construct. Steel Res. 106, 311 321. Denoon, R.O., 2000. Designing for Serviceability Accelerations in Buildings. PhD Thesis. University of Queensland, Queensland, Australia. Derham, C.G., Kelly, J.M., Thomas, A.G., 1985. Nonliner natural rubber bearings for seismic isolation. Nuclear Eng. Design. 84 (3), 417 428. Deulkar, W., Modhera, C., Patil, H., 2010. Buckling restrained braces for vibration control of building structures. Int. J. Recent Res. Appl. Stud. 4, 363 372. DG/TJ08-15, 2004. Code for Design of Steel-Concrete Hybrid Structures for High-rise Buildings. Chinese Standard. Dicleli, M., Buddaram, S., 2007. Comprehensive evaluation of equivalent linear analysis method for seismic-isolated structures represented by SDOF systems. Eng. Struct. 29 (8), 1653 1663. Dima, A., Stef˘ ¸ anescu, B., 2003. A seismic base isolation device. In: Proceedings of the 10th International Conference on Metal Structures, Timi¸soara, Romania, pp. 346 351. Di Matteo, A., Lo Iacono, F., Navarra, G., Pirrotta, A., 2014. Optimal tuning of tuned liquid column damper systems in random vibration by means of an approximate formulation. Meccanica 50 (3), 795 808. DIN, 2005. DIN 1055-4:2005-03 Einwirkungen auf Tragwerke Teil 4: Windlasten. Diotallevi, P.P., Landi, L., Lucchi, S., 2014. A direct method for the design of viscous dampers to be inserted in existing buildings, In: Tenth U.S. National Conference on Earthquake Engineering Frontiers of Earthquake Engineering, Anchorage, Alaska, July 21 25. Di Paola, M., Navarra, G., 2008. Passive control of linear structures equipped with nonlinear viscous dampers and amplification mechanisms. In: Proc. of 52th Giornata di Aggiornamento CIAS Centro Internazionale di Aggiornamento SperimentaleScientifico, Palermo, Italy. DIS, 2017. Viscous Wall Damper: Modeling Guide. Dynamic Isolation System, Nevada, USA. Dogangun, A., 2009. Analysis and Design of the Concrete Structures: TDY 2007. Birsen Press, Istanbul. Dolce, M., Marnetto, R., 2000. Passive seismic devices based on shape memory alloys. In: Proceedings of the 12th World Conference on Earthquake Engineering, Auckland, New Zealand.

References

Dolce M., Filardi B., Marnetto R., Nigro D., 1996. Experimental tests and applications of a new biaxial elasto-plastic device for the passive control of structures. In: Fourth world congress on joint sealants and bearing systems for concrete structures, ACI SP-164, Sacramento, California. Domizio, M., Ambrosini, D., Curadelli, O., 2015. Performance of tuned mass damper against structural collapse due to near fault earthquakes. J. Sound Vibrat. 336 (3), 32 45. Driver, R.G., Kulak, G.L., Kennedy, D.J.L., Elwi, A.E., 1997. Seismic behavior of steel plate shear walls. In: Structural Engineering Rep. No. 215, Dept. of Civil Engineering, Univ. of Alberta, Edmonton, Canada. Du, H., Boffa, J., Zhang, N., 2006. Active seismic response control of tall buildings based on reduced order model. In: Proceedings of the 2006 American Control Conference Minneapolis, Minnesota, USA, June 14 16. Dyke, S.J., Spencer Jr., B.F., Quast, P., Sain, M.K., Kaspari Jr., D.C., Soong, T.T., 1996a. Acceleration feedback control of MDOF structures. J. Eng. Mech. ASCE 120 (2), 266 283. Dyke, S.J., Spencer Jr., B.F., Sain, M.K., Carlson, J.D., 1996b. Marketing and control of magneto-rheological dampers for seismic response reduction. Smart Mater. Struct. 5, 565 575. Dyke, S.J., Yi, F., Frech, S., Carlson, J.D., 1999. Application of magnetorheological dampers to seismically excited structures. In: International Modal Analysis Conference, IMAC XVII, Florida, USA. EERI, 2012. Reconnaissance Team: Performance of engineered structures in the Mw 9.0 Tohoku, Japan, Earthquake of March 11, 2011. EERI Special Earthquake Report, Earthquake Engineering Research Institute, Oakland, CA. Ekwueme, C.G., Hart, G.C., Ozegbe, K., Brandow, G., 2010. Using Viscous Dampers on A 42-Story High-Rise Building. 2010 Annual Meeting, Los Angeles Tall Buildings Structural Design Council, At Los Angeles, CA. Elbakheit, A.R., 2007. Enhanced Integration of Photovoltaic and Wind Turbines Into Building Design. PhD thesis, the British Library. Elbakheit, A.R., 2014. Factors enhancing aerofoil wings for wind energy harnessing in buildings. Build. Serv. Eng. Res. Technol. 35 (4), 417 437. Elgaaly, M., 1998. Thin steel plate shear walls behavior and analysis. Thin-Wall. Struct. 32, 151 180. Elgaaly, M., Caccese, V., Du, C., 1990. Steel plate shear walls post-buckling under cyclic loads. In: Proc. of the Fourth U.S. Nat. Conf. on Earthquake Engineering, Earthquake Eng. Res. Inst. (EERI), E1 Cerrito, Calif. Elias, S., Matsagar, V., 2014. Distributed multiple tuned mass dampers for wind vibration response control of high-rise building. J. Eng. 2014, 1 11. Ellis, B.R., 1996. Full-scale measurements of the dynamic characteristics of buildings in the UK. J. Wind Eng. Ind. Aerodyn. 59, 35 382. EN, 2003. EN 1337-10: Structural bearings. Inspection and maintenance. BSI. EN ISO 9001, 2008. Quality Management Systems Requirements. CEN. Erlandsson, M., Borg, M., 2003. Generic LCA-methodology applicable for buildings, constructions and operation services—today practice and development needs. Build. Environ. 38, 919 938. Erwin, S., Kijewski-Correa, T., Yoon, S.W., 2007. Full-scale Verification of Dynamic Properties from Short Duration Records. 2007 Structures Congress: New Horizons and Better Practices, Structural Engineering Institute (SEI), Reston, VA.

1025

1026

References

ESDU, 2012. ESDU 83009, Damping of Structures Part 1: Tall Buildings. Engineering Science Data Unit International PLC, London, UK. ES ISO, 2001. ISO 3010: 2001 Basis for Design of Structures Seismic Actions on Structures. International Organization for Standardization. ES ISO, 2009. ISO 4354: 2009 Wind on Structures. International Organization for Standardization. Eskildsen, P.E., 1965. The World Trade Center Wind Effects No. 1. Oregon Research Inst, Eugene, Oregon. Eskildsen, P.E., 1966. The World Trade Center Wind Effects No. s2. Oregon Research Inst, Eugene, Oregon. Facioni, R.J., Kwok, K.C.S., Samali, B., 1995. Wind tunnel investigation of active vibration control of tall buildings. J. Wind Eng. Ind. Aerodyn. 54 55, 397 412. Falcon, K.C., Store, B.J., Simcock, W.D., Andrew, C., 1967. Optimization of vibration absorbers: a graphical method for use on idealized systems with restricted damping. J. Mech. Eng. Sci. 9, 374 381. Falconer, D.W., 1981. Classification of Tall Building Systems. Master of Science Thesis, Lehigh University. Faltinsen, O.M., Timoka, A.N., 2009. Sloshing. Cambridge University Press, Cambridge. Faltinsen, O.M., Firoozkoohi, R., Timoka, A.N., 2011. Analytical modeling of liquid sloshing in a two-dimensional rectangular tank with a slat screen. J. Eng. Math. 70 (1), 93 109. Farghaly, A., Ahmed, M.S., 2012. Optimum design of TMD System for tall buildings. J. Eng. 2012, 1 13. Faridani, H.M., 2015. Replacement beam methods in analysis of tall building structural systems. PhD dissertation, Politecnico di Milano, Italy. Faridani, H.M., Capsoni, A., 2016a. Investigation of the effects of viscous damping mechanisms on structural characteristics in coupled shear walls. Eng. Struct. 116 (2016), 121 139. Faridani, H.M., Capsoni, A., 2016b. A modified replacement beam for analyzing building structures with damping systems. Struct. Eng. Mech. 58 (5), 905 929. Farshidianfar, A., Oliazadeh, P., 2009. Closed form optimal solution of a tuned liquid column damper responding to earthquake. Int. J. Civil Environ. Struct. Construct. Arch. Eng. 3 (11), 457 462. Farshidianfar, A., Soheili, S., 2013. Ant colony optimization of tuned mass dampers for earthquake oscillations of high-rise structures including soil-structure interaction. Soil Dyn. Earthquake Eng. 51, 14 22. Fava, J., Denison, R., Jones, B., Curran, M., Vigon, B., Selke, S., et al., 1991. A Technical Framework for Life-Cycle Assessment. SETAC Press, Pensacola, FL. Feese, C.A., 2013. Assessment of Seismic Damage of Buildings and Related Environmental Impacts. Master Thesis, Michigan Technological University. FEMA, 1992a. NEHRP Handbook for the Seismic Evaluation of Existing Buildings, FEMA 178 report, Building Seismic Safety Council, Federal Emergency Management Agency, Washington, DC. FEMA, 1992b. NEHRP Handbook of Techniques for the Seismic Rehabilitation of Existing Buildings, FEMA 172 report, Building Seismic Safety Council, Federal Emergency Management Agency, Washington, DC.

References

FEMA, 1997. NEHRP Guidelines for the Seismic Rehabilitation of Buildings, Federal Emergency Management Agency, FEMA 273 report, Washington, DC. FEMA, 1998a. Handbook for the Seismic Evaluation of Buildings, Federal Emergency Management Agency, FEMA 310 report, Washington, DC. FEMA, 1998b. Planning for seismic rehabilitation: societal issues. FEMA 275, prepared by VSP Associates for the Building Seismic Safety Council and Federal Emergency Management Agency, Washington, DC. FEMA, 2000a. Recommended Seismic Design Criteria for New Steel Moment-Frame Buildings, Federal Emergency Management Agency, FEMA 350 report, Washington, DC. FEMA, 2000b. Recommended Seismic Evaluation and Upgrade Criteria for Existing Welded Steel Moment-Frame Buildings, Federal Emergency Management Agency, FEMA 351 report, Washington, DC. FEMA, 2000c. Recommended Postearthquake Evaluation and Repair Criteria for Welded Steel Moment-Frame Buildings, Federal Emergency Management Agency, FEMA 352 report, Washington, DC. FEMA, 2000d. Prestandard and Commentary for the Seismic Rehabilitation of Buildings, Federal Emergency Management Agency, FEMA 356 report, Washington, DC. FEMA, 2001. NEHRP Recommended Provision for Seismic Regulations for New Buildings and other Structures. Federal Emergency Management Agency, FEMA-368 report. FEMA, 2002. Hazard Risk Assessment Program Hazus®: Hazus-MH 3.1. Federal Emergency Management Agency. FEMA, 2003. FEMA 450: NEHRP Recommended Provisions for Seismic Regulations for New Buildings and Other Structures. Building Seismic Safety Council of the National Institute of Building Sciences, Washington, D.C. FEMA, 2004. FEMA389: Primer for Design Professionals: Communicating With Owners and Managers of New Buildings on Earthquake Risk. Federal Emergency Management Agency. FEMA, 2005. FEMA 400: NEHRP Improvement of Nonlinear Static Seismic Analysis Procedures. Federal Emergency Management Agency ASCE, Washington, DC. FEMA, 2006a. Risk Management Series, Designing for Earthquakes: A Manual for Architects. Federal Emergency Management Agency, Washington, DC. FEMA, 2006b. FEMA 451: NEHRP Recommended Provisions: Design Examples. Procedures. Federal Emergency Management Agency ASCE, Washington, DC. FEMA, 2009. FEMA P695: Quantification of Building Seismic Performance Factors. Federal Emergency Management Agency, Washington, DC. FEMA, 2011. FEMA E-74: Reducing the Risks of Nonstructural Earthquake Damage A Practical Guide. Federal Emergency Management Agency, Washington, DC. FEMA, 2012. FEMA P-58: Seismic Performance Assessment of Buildings. Federal Emergency Management Agency, Washington, DC. FEMA, 2015. Hazus MH 2.1: Earthquake Model Technical Manual. Federal Emergency Management Agency, Washington, DC. Feng, D., etc. 2012. A new design procedure for seismically isolated buildings based on seismic isolation codes worldwide, 15th WCEE, Paper No. 0435. Ferritto, J.M., 1984. Economics of seismic design for new buildings. J. Struct. Eng. 110 (No. 12). Filiatrault, A., Tremblay, R., Karr, R., 2000. Performance evaluation of friction spring seismic damper. ASCE J. Struct. Eng. 126 (4), 491. 499.

1027

1028

References

Filiatrault, A., Christopoulos, C., Stearns, C., 2002. Guidelines, Specifications, and Seismic Performance Characterization of Nonstructural Building Components and Equipment. PEER Report 2002/05. Fisco, N.R., Adeli, H., 2011a. Smart structures: Part I active and semi-active control. Sci. Iranica 18 (3), 275 284. Fisco, N.R., Adeli, H., 2011b. Smart structures: Part II hybrid control systems and control strategies. Sci. Iranica 18 (3), 285 295. Frahm, H., 1911. Device for Damping of Bodies. US Patent No: 989,958. Franco-Villafan˜e, J.A., Flores, J., Mateos, J.L., Mendez-Sanchez, R.A., Novaro, O., Seligman, T.H., 2011. Novel doorways and resonances in large-scale classical systems. Lett. J. Expl. Front. Phys. EPL 94 (30005), (5pp). Frandsen, J.B., 2005. Numerical predictions of tuned liquid damper systems. J. Fluids Struct. 20 (3), 309 329. Fritz, W.P., Jones, N.P., Igusa, T., 2009. Predictive models for the median and variability of building period and damping. J. Struct. Eng. 135 (5), 576. 576 586. Fu, T.S., Johnson, A., 2014. Structural health monitoring with a distributed mass damper system. Struct. Control Health Monit. 21 (2), 189 204. Fu, T.S., Zhang, R., 2016. Integrating double-skin Fac¸ades and mass dampers for structural safety and energy efficiency. J. Arch. Eng. 1 12. Fujino, Y., Sun, L.M., 1993. Vibration control by multiple tuned liquid dampers (MTLDs). J. Struct. Eng. 119 (12), 3482 3502. Fujino, Y., Sun, L., Pacheco, B., Chaiseri, P., 1992. Tuned liquid damper (TLD) for suppressing horizontal motion of structures. J. Eng. Mech. 118 (10), 2017 2030. Fujita, K., Takewaki, I., 2011a. Earthquake response bound analysis of uncertain baseisolated buildings for robustness evaluation. J. Struct. Construct. Eng. 666, 1453 1460 (in Japanese). Fujita, K., Takewaki, I., 2011b. An efficient methodology for robustness evaluation by advanced interval analysis using updated second-order Taylor series expansion. Eng. Struct. 33 (12), 3299 3310. Fujita, K., Takewaki, I., 2012. Robustness evaluation on earthquake response of baseisolated buildings with uncertain structural properties under long-period ground motions. Arch. J. 1 (1), 46 59. Fujitani, H., Sodeyama, H., Hata, K., Komatsu, Y., Iwata, N., Sunakoda, K., et al., 2002. Dynamic performance evaluation of 200kN Magneto-pheological Damper. Tech. Note Natl. Inst. Land Infrastruct. Manage. 41, 349 356. Fuller, C.R., Elliott, S.J., Nelson, P.A., 1996. Active Control of Vibration. Academic Press, London. Gamota, D.R., Filisko, F.E., 1991. Dynamic mechanical studies of electrorheological materials: moderate frequencies. J. Rheol. 35, 399 425. Gao, H., Kwok, K.C.S., Samali, B., 1997. Optimisation of tuned liquid column dampers. Eng. Struct. 19 (6), 476 486. Gao, H., Kwok, K.C.S., Samali, B., 1999. Characteristics of multiple tuned liquid column dampers in suppressing structural vibration. Eng. Struct. 21, 316 331. Garevski, M., 2013. Earthquakes and Health Monitoring of Civil Structures. Springer, Netherlands. Gatscher, J.A., Bachman, R.E., 2012. Elements of 2012 IBC / ASCE 7-10 Nonstructural Seismic Provisions: Bridging the Implementation Gap. 15 WCEE, LISBOA 2012.

References

Gavin, H., Alhan, C., Oka, N., 2003. Fault tolerance of semiactive seismic isolation. J. Struct. Eng. 129 (7), 922 932. GB 50223, 2008. China Ministry of Construction (CMC). Standard for Classification of Seismic Protection of Buildings Constructions. China Architecture & Building Press, Beijing, China. GB 50009, 2012. China Ministry of Construction (CMC). Load Code for the Design of Building Structures. China Architecture & Building Press, Beijing, China. GB 50982-2014, Technical code for monitoring of building and bridge structures (in Chinese). GB 50010, 2010. China Ministry of Construction (CMC). Code for Design of Concrete Structures. China Architecture & Building Press, Beijing, China. GB 50011, 2010. China Ministry of Construction (CMC). Code for Seismic Design of Buildings. China Architecture & Building Press, Beijing, China. GB 50017, 2010. China Ministry of Construction (CMC). Code for Design of Steel Structures. China Architecture & Building Press, Beijing, China. GB 50936, 2014. China Ministry of Construction (CMC). Technical Code for Concrete Filled Steel Tubular Structures. China Architecture & Building Press, Beijing, China. Gerges, R., Vickery, B.J., 2005. Optimum design of pendulum-type tuned mass dampers. Struct. Design Tall Spec. Build. 14 (4), 353 368. Gerges, R.R., Vickery, B.J., 2003. Parametric experimental study of wire rope spring tuned mass dampers. J. Wind Eng. Ind. Aerodyn. 91 (12-15), 1363 1385. Gheidi, A., Mirtaheri, M., Zandi, A.P., Alanjari, P., 2009. Effect of filler material on local and global behaviour of buckling-restrained braces. Struct. Design Tall Spec. Build. 1 11. Gluck, N., Reinhorn, A.M., Gluck, J., Levy, R., 1996. Design of supplemental dampers for control of structures. J. Struct. Eng. 122 (12), 1394 1399. Goel, R.K., Chopra, A.K., 1997. Vibration Properties of Buildings during Earthquakes, Rep. to be Published, Earthquake Eng. Res. Ctr.• Univ. of California. Berkeley, Richmond, Calif. Gong, S., Zhou, Y., 2016. Experimental study and numerical simulation on a new type of viscoelastic damper with strong nonlinear characteristics. Struct. Control Health Monit. Available from: https://doi.org/10.1002/stc.1897. Gordaninejad, F., Saiidi, M., Hansen, B.C., Ericksen, E.O., Chang, F.-K., 2002. Magnetorheological fluid dampers for control of bridges. J. Intell. Mater. Syst. Struct. 13 (2 3). GOST R. 53778, 2010. Building and structures, technical inspections and monitoring regulations. Gray, M.G., Christopoulos, C., Packer, J.A., Lignos, D.G., 2012. Development, Validation and Modeling of the new Cast Steel Yielding Brace System. Proceedings, Structures Congress, March 29 31, 2012 | Chicago, Illinois, USA. Gray, M.G., Christopoulos, C., Packer, J.A., 2014. Cast steel yielding brace system for concentrically braced frames: concept development and experimental validations. Struct. Eng. 140 (4), 04013095. Greenblatt, D., Wygnanski, I., 2000. The control of flow separation by period excitation. Prog. Aerospace Sci. 36, 487 545. Grigorian, C.E., Yang, T., Popov, E.P., 1993. Slotted bolted connection energy dissipators. Earthquake Spectra 9 (3), 491 504. Gu, M., Quan, Y., 2004. Across-wind loads of typical tall buildings. J. Wind Eng. Ind. Aerodyn. 92, 1147 1165.

1029

1030

References

Gunturi, S.K.V., Shah, H.S., 1993. Mapping Structural Damage to Monetary Damage. Structural Engineering in Natural Hazards Mitigation: Proc. ASCE Structures Congress, Irvine, CA, 1331-1336. Guo, Y.L., Kareem, A., Ni, Y.Q., Liao, W.Y., 2012. Performance evaluation of Canton Tower under winds based on full-scale data. J. Wind Eng. Ind. Aerodyn. 104-106, 116 128. Guo, J.W.W., Christopoulos, C., 2013. A procedure for generating performance spectra for structures equipped with passive supplemental dampers. Earthquake Eng. Struct. Dyn. 42, 1321 1338. Hadi, M.N.S., Arfiadi, Y., 1998. Optimum design of absorber for MDOF structures. J. Struct. Div. 124, 1272 1280. Halabian, A.M., Torki, M., 2011. Numerical studies on the application of tuned liquid dampers with screens to control seismic response of structures. Struct. Design Tall Spec. Build. 20 (2), 121 150. Hall, J.F., 2005. Problems encountered from the use (or misuse) of Raleigh damping. Earthquake Eng. Struct. Dyn. 35 (5), 525 545. Hall, J.F., Ryan, K.L., 2000. Isolated buildings and the 1997 UBC near-source factors. Earthquake Spectra 16 (2), 393 411. Halvarson, R., Isyumov, N., 1986. Comparison of Predicted and Measured Dynamic Behaviour of Allied Bank Plaza. Building Motion in Wind, American Society of Civil Engineers, New York, pp. 23 41. Han, S., Klein, G., Grossman, G., 2012. Cost-effective solution: Controlling perception of win-induced motion in slender buildings. Structure Magazine, June. Han, Y.-L., Li, Q.S., Li, A.-Q., Leung, A.Y.T., Lin, P.-H., 2003. Structural vibration control by shape memory alloy damper. Earthquake Eng. Struct. Dyn. 32 (3), 483 494. Hanagan, L.M., Murray, T.M., 1997. Active control approach for reducing floor vibration. J. Struct. Eng. 123 (11), 1497 1505. Hanson, R.D., Aiken, I.D., Nims, D.K., Richter, P.J., Bachman, R.E., 1993. State-of-the-art and State-of-the-practice in Seismic Energy Dissipation. Haskett, T., Breukelman, B., Robinson, J., Kottelenberg, J., 2003. Tuned Mass Dampers under excessive structural excitation, Extreme Loading Conference, Toronto, Canada. Hasmimoto, T., Fuijta, K., Tsuji, M., Takewai, I., 2015. Innovative base-isolated building with large mass-ratio TMD at basement for greater earthquake resilience. Future Cities Environ. 1 (9). Hasselman, T.K., Wiggins J.H., 1982. Earthquake damage to High-rise Buildings as a function of Interstory Drift. In: Proc. 3rd Int. Earthquake Microzonation Conf., Seattle, WA, 883 894. Haviland, R., 1976. A study of the uncertainties in the fundamental translational periods and damping values for real buildings. Research Rep. No. 5, Massachusetts Institute of Technology, Cambridge, MA. Hazaveh, N.K., Pampanin, S., Rodgers, G.W., Chase, J.G., 2014. Novel semi-active viscous damping device for reshaping structural response. In: Conference: 6WCSCM (Sixth World Conference of the International Association for Structural Control and Monitoring). Hazaveh, N.K., Rodgers, G.W., Pampanin, S., Chase, J.G., 2016. Damping reduction factors and code-based design equation for structures using semi-active viscous dampers. Earthquake Eng. Struct. Dyn. Available from: https://doi.org/10.1002/eqe.2782.

References

He, Y.C., Li, Q., 2014. Dynamic responses of a 492-m-high tall building with active tuned mass damping system during a typhoon. Struct. Control Health Monit. 21 (5), 705 720. Hedemann, J., Ko¨nig, U., 2007. Technical Documentation of the Ecoinvent Database. Final Report Ecoinvent data v2.0, No. 4. Swiss Centre for Life Cycle Inventories, Du¨bendorf. Heo, J.-S., Lee, S.-K., Park, E., Lee, S.-H., Min, K.-W., Kim, J.J., et al., 2009. Performance test of a tuned liquid mass damper for reducing bidirectional responses of building structures. Struct. Design Tall Spec. Build. 18 (7), 789 805. Herve´ Poh’sie´, G., Chisari, C., Rinaldin, G., Fragiacomo, M., Amadio, C., Ceccotti, A., 2015. Application of a translational tuned mass damper designed by means of genetic algorithms on a multistory cross-laminated timber building. J. Struct. Eng. 142 (4). Higashino, M., Okamoto, S. (Eds.), 2006. Response Control and Seismic Isolation of Buildings. Taylor & Francis, New York. Hino, J., Yoshitomi, S., Tsuji, M., Takewaki, I., 2008. Bound of aspect ratio of baseisolated buildings considering nonlinear tensile behavior of rubber bearing. Struct. Eng. Mech. 30 (3), 351 368. Hitaka, T., Matsui, C., 2003. Experimental study on steel sear wall with slits. J. Struct. Eng. 129 (5), 586 595. Hitaka, T., Ito, M., Murata, Y., Nakashima, M., 2009. Seismic behavior of steel shear plates stiffened by wood panels. In: Proceedings of STESSA 2009 Behaviour of Steel Structures in Seismic Areas, Philadelphia, USA, 623 628. Hitchcock, P.A., Kwok, K.C.S., Watkins, R.D., Samali, B., 1997. “Characteristics of liquid column vibration absorbers (LCVA) I”. Eng. Struct. 19 (2), 126 134. HKBD, 2004. Code of practice on Wind Effects in Hong Kong. Buildings Department, The Government of the Hong Kong Special Administrative Region. HKBD, 2011. Code of Practice for Structural use of Steel. Buildings Department, The Government of the Hong Kong Special Administrative Region. HKBD, 2013. Code of Practice for Structural use of Concrete. Buildings Department, The Government of the Hong Kong Special Administrative Region. Ho, G.W.M., 2016. The evolution of outrigger system in tall buildings. Int. J. High-Rise Build. 5 (1), 21 30. Hoang, N., Fujino, Y., Warnitchai, P., 2008. Optimal tuned mass damper for seismic applications and practical design formulas. Eng. Struct. 30, 707 715. Hori, N., Seto, K., 2000. Vibration control of flexible space structures based on reducedorder modeling method and filtered LQ control theory. JSME Int. J. Ser. C Mech. Syst. Mach. Elem. Manuf. 43 (3), 697 703. Horiuchi, T., Konno, T., 2001. A new method for compensating actuator delay in real-time hybrid experiments. R. Soc. Lond. 359, 1893 1909. Hou, Z.N., Feng, Z.M., Hu, H.G., Wu, G.B., 2010. Experimental study on performance characteristics of magnetorheological damper. Appl. Mech. Mater. 37 38, 439 443. Hunt, S.J., 2002. Semi-active smart-dampers and resettable actuators for multi-level seismic hazard mitigation of steel moment resisting frames. Masters Thesis, Mechanical Engineering, University of Canterbury, Christchurch. Huo, L., Li, H., 2011. Seismic Response Reduction of Eccentric Structures Using Liquid Dampers, Vibration Analysis and Control New Trends and Developments, Dr. Francisco Beltran-Carbajal (Ed.).

1031

1032

References

Huo, L., Shen, W., Li, H., Zhang, Y., 2013. Optimal design of liquid dampers for structural vibration control based on GA and HN norm. Math. Prob. Eng. 2013, 1 10. Hussain, S., Benschoten, P.V., Satari, M.A., Lin, S., 2006. Buckling Restrained Braced Frames (BRBF) structures: analysis, design and approvals issues. In: Proceedings of the 75th SEAOC Annual Convention. Long Beach: Coffman Engineering. Hwang, A., 1998. Viscous Dampers: Practical Application Issues for the Structural Engineer. Master of Science Thesis, Massachusetts Institute of Technology. Hwang, J.S., Chiou, J.M., 1996. An equivalent linear model of lead-rubber seismic isolation bearings. Eng. Struct. 18 (7), 528 536. Hwang, J.S., Huang, Y.N., Yi, S.L., Ho, S.Y., 2008. Design formulations for supplemental viscous dampers to building structures. J. Struct. Eng. 134 (1), 22 31. Hwang, J.S., Lin, W.C., Wu, N.-J., 2013. Comparison of distribution methods for viscous damping coefficients to buildings. Struct. Infrastruct. Eng. 9 (1), 28 41. IBC, 2009. International Building Code 2009. International Code Council, Inc. Proceedings of the 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada. IBC, 2012. International Building Code 2012. International Code Council, Inc. IBC, 2015. International Building Code 2015. International Code Council, Inc. IBC, 2018. International Building Code 2018. International Code Council, Inc. Ibrahim, R.A., 2005. Liquid Sloshing Dynamics: Theory and Applications. Cambridge University Press, Cambridge. ICC, 2003. International Building Code. International Code Council, Building Officials and Code Administrators International Inc., Country Club Hills, IL., International Conference of Building Officials, Whittier, CA, Southern Building Code Congress International, Inc., Birmingham, Alabama. ICC, 2009. International Building Code. Falls Church, VA: International Code Council. IDEMAT, 2001. Database 2001. Faculty of Industrial Design Engineering of Delft University of Technology, The Netherlands. Iemura, H., Igarashi, A., Tanaka, H., 2002. Substructure hybrid shake table test using digitally filtered displacement compensation control scheme. In: Proceedings of the 11th Japan Earthquake Engineering Symposium, No. 304 (in Japanese). Iemura, H., Igarashi, A., Toyooka, A., Suzuki, Y., 2003. Dynamic loading experiment of passive and semiactive dampers using the inertia-force-driven loading system. 2, 45 50. In: Proceedings of the Third World Conference on Structural Control, Como, Italy. Iemura, H., Igarashi, A., Kalantari, A., 2006. Experimental verification and numerical studies of an autonomous semi-active control strategy. Struct. Control Health Monit. 13 (1), 301 323. Igarashi, A., 1994. On-line computer controlled testing of stiff multi-DoF structural systems under simulated seismic loads. Ph.D. Thesis, University of California, San Diego. Igarashi, A., Seible, F., Hegemier, G.A., 1993. Development of the pseudodynamic technique for testing a full-scale 5-story shear wall structure. U.S.-Japan Seminar on Development and Future Dimension of Structural Testing Techniques. Igusa, T., Xu, K., 1994. Vibration control using multiple tuned mass damper. J. Sound Vibrat. 175, 491 503. Ikeda, Y., 2009. Active and semi-active vibration control of buildings in Japan—Practical applications and verification. Struct. Control Health Monit. 16 (7 8), 703 723. IMC, 2012. International Mechanical Code 2012. International Code Council, Inc.

References

Inaudi, J.A., Kelly, J.M., To, C.W.S., 1993. Statistical Linearization Method in the Preliminary Design of Structures with Energy Dissipation Devices. Proc. of Seminar on Seismic Isolation, Passive Energy Dissipation and Active Control, Applied Technological Council, ATC 17-1, Vol. 2, pp. 502 520. Indirli, M., Castellano, M.G., 2008. Shape memory alloy devices for the structural improvement of masonry heritage structures. Int. J. Arch. Heritage 2, 93 119. Infanti, S., Robinson, J., Smith R., 2008. Viscous dampers for high-rise buildings. The 14 th World Conference on Earthquake Engineering, October 12 17, Beijing, China. Inman, D.G., 2006. Vibration with Control. John Wiley & Sons Ltd, The Atrium, Southern Gate, Chichester, West Sussex PO19 8SQ, England. Inoue, K., Kuwahara, S., 1998. Optimum strength ratio of hysteretic damper. Earthquake Eng. Struct. Dyn. 27 (6), 577 588. International Conference of Building Officials (ICBO), 1997. UBC: Earthquake regulations for seismic isolated structures. In: Uniform building code, USA. Irandegani, M.A., Narmashiri, K., Rahgozar, R., Rasoolina, M., 2014. Aluminum panel yielding dampers: application in steel structures. Int. Res. J. Appl. Basic Sci. 8 (9), 1459 1470. Irikura, K., Kamae, K., Kawabe, H., 2004. Importance of prediction of long period ground motion during large earthquakes. Annual Conference of the Seismological Society of Japan, Poster session (in Japanese). Irwin, P., Kilpatrick, J., Robinson, J., Frisque, A., 2008a. Wind and tall buildings: negatives and positives. Struct. Design Tall Spec. Build. 17, 915 928. Irwin, P.A., Kilpatrick, J., Frisque, A., 2008b. Friend or Foe, Wind at Height. Proceedings of CTBUH 8th World Congress, Dubai. Irwin, P.A., 2008. Bluff body aerodynamics in wind engineering. J. Wind Eng. Ind. Aerodyn. 96, 701 712. Irwin, P.A., Breukelman, B., 2001. Recent applications of damping systems for wind response. In: Conference Proceedings Of the 6th CTBUH World Congress, Melbourne. Isermann, R., Mu¨nchhoff, M., 2011. Identification of Dynamic Systems: An Introduction with Applications. Springer, New York. Ishikawa, K., Okuma, K., Shimada, A., Nakamura, H., Masaki, N., 2004. JSSI manual for building passive control technology part-5 performance and quality control of viscoelastic dampers. In: 13 th World Conference on Earthquake Engineering, Vancouver, BC, Canada August 1 6. ISIS, 2001. Guidelines for Structural Health Monitoring—Design Manual No. 2. ISIS Canada—the Canadian Network of Centres of Excellence on Intelligent Sensing for Innovative Structures, University of Manitoba, Winnipeg, Manitoba, Canada. Iskhakov, I., Ribabok, Y., 2007. Modern trends in base isolation applications for seismic protection of modern buildings. In: Proceedings of the 10th International Conference on Studies, Repairs and Maintenance of Heritage Architecture, Prague, Czech Republic, pp. 623 632. Islam, B., Ahsan, R., 2012. Optimization of tuned mass damper parameters using evolutionary operation algorithm. In: 15th World Conference in Earthquake Engineering (WCEE), Lisbon, Portugal. Islam, A.S., Ahmad, S.I., Jameel, M., Zamin, M.J., 2012. Seismic base isolation for buildings in regions of low to moderate seismicity: practical alternative design. Pract. Period. Struct. Design Construct. 17 (1), 13 20.

1033

1034

References

Islam, A.B.M.S., Jameel, M., Uddin, M.A., Ahmad., S.I., 2011. Simplified design guidelines for seismic base isolation in multi-storey buildings for Bangladesh National Building Code (BNBC). Int. J. Phys. Sci. 6 (23), 5467 5486. ISO 10137, 2007. ET ISO 10137: Bases for design of structures serviceability of buildings and walkways against vibrations. International Organization for Standardization. ISO 14040, 1997. Environmental Management Life Cycle Assessment Principles and Framework. International Standard. International Organization for Standardization. ISO 14044, 2006. Environmental Management Life Cycle Assessment Requirements and Guidelines. International Standard. International Organization for Standardization. ISO 22762-1, 2010. Elastomeric seismic-protection isolators Part 1: Test methods. ISO. ISO 6897, 1984. Guidelines for the evaluation of the response of occupants of fixed structures, to low-frequency horizontal motion (0.063 to 1 Hz). International Organization for Standardization. ISO 9001, 2008. Quality management systems Requirements. ISO. ISO 9001, 2015. Quality Management Systems Requirements. International Standard Organization, Geneva, Swizterland. ISO/TC 178 N 484, 2006. Lifts (elevators) Study into the use of lifts for evacuation during an emergency. 10 April 2006. Isyumov, N., 1995. Motion Perception, Tolerance and Mitigation. 5th World Congress of the Council on Tall Buildings and Urban Habitat 14 19, 1995. Amsterdam, Netherlands, May. Isyumov, N., Case, P., Hasan, A., Tait, M., 2010. Monitoring of Tall Buildings to Assist the Design of Supplementary Damping Systems. Structures Congress 2010, Orlando, Florida. Iwanami, K., Seto, K., 1984. Optimum design of dual tuned mass dampers and their effectiveness. Jpn Soc. Mech. Eng. 50 (1), 44 52. Iwata, M., 2004. Applications-design of buckling restrained braces in Japan. Jackson, M., Scott, D.M., 2010. Increasing efficiency in tall buildings by damping. Structures Congress. May 12 15, Orlando, Florida, USA. Jacobsen, L.S., 1930. Steady force vibration as influenced by damping. Trans. Am. Soc. Mech. Eng. 52 (15), 169 181. Jagadish, G.K., Jangid, R.S., 2009. Semi-active MR damper for seismic control of structures. Bull. NZ Soc. Earthquake Eng. 42 (3), 157 166. Jain, S.K., Thakkar, S.K., 2004. Application of base isolation for flexible buildings. In: Proceedings of the 13th World Conference on Earthquake Engineering, Vancouver, Canada. Jangid, R.S., 1995. Dynamic characteristic of structures with multiple tuned mass dampers. Struct. Eng. Mech. 3, 497 509. Jangid, R.S., 1999. Optimum multiple tuned mass dampers for base-excited undamped system. Earthquake Eng. Struct. Dyn. 28 (9), 1041 1049. Jara, M., Casas, J.R., 2006. A direct displacement-based method for the seismic design of design of bridges on bilinear isolation devices. Eng. Struct. 28 (6), 869 879. JCC, 2006. California Trial Court Facility Standards. Judicial Council of California San Francisco, CA. Jeary, A.P., 1986. Damping in buildings a mechanism and a predictor. J. Earthquake Eng. Struct. Dyn. 14, 733 750. Jeary, A.P., 1996. The description and measurement of nonlinear damping in structures. J. Wind Eng. Ind. Aerodyn. 59, 103 104.

References

Jeong, H.K., Goncalves, F.D., Ahmadian, M., 2004. Investigation of the response time of magnetorheological fluid dampers. In: Proc. Smart Structures and Materials: Damping and Isolation, 29 th July, Vol. 5386, 63-71, San Diego, USA. JGJ 6, 2011. China National Building Code. Technical Code for Tall Building Raft Foundations and Box Foundations. JGJ 99, 1998. China National Building Code. Technical Specification for steel structures of tall buildings. JGJ 138, 2001. China National Building Code. Technical Specification for Steel Reinforced Concrete Composite Structures. JGJ3, 2010. China National Building Code. Technical Specification for Concrete Structures of Tall Building. JGJ297, 2013. Technical Specification for Building with Energy Dissipation Devices. China National Building Code. JG/T209, 2017. Dampers for Vibration Energy Dissipation of Buildings. China National Building Code. Jia, J., 2015. Modern Earthquake Engineering: Offshore and Land-based Structures. Springer, Germany. Jiang, Q., Lu, X., Guan, H., Ye, X., 2014. Shaking table model test and FE analysis of a reinforced concrete mega-frame structure with tuned mass dampers. Struct. Design Tall Spec. Build. 23 (18), 1426 1442. Jiao, Y., Kishiki, S., Yamada, S., Ene, D., Konishi, Y., Hoashi, Y., et al., 2015. Low cyclic fatigue and hysteretic behavior of U-shaped steel dampers for seismically isolated buildings under dynamic cyclic loadings. Earthquake Eng. Struct. Dyn. 44 (10), 1523 1538. Jin, S., Ou, J., Richard Liew, J.Y., 2016. Stability of buckling-restrained steel plate shear walls with inclined-slots: theoretical analysis and design recommendations 117, 13 23. Jin, H.J., Li, G.Q., Lu, Y., Sun, F.F., He, Y.M., 2014. Buckling-Restrained Steel Plate Shear Walls Demands on buckling restrainers. In: Proceedings of EUROSTEEL 2014 the 7th European Conference on Steel and Composite Structures, on stick, Napoli, Italia. Jolliet, O., Margni, M., Charles, R., Humbert, S., Payet, J., Rebitzer, G., et al., 2003. IMPACT 2002 1 : a new life cycle impact assessment methodology. Int. J. Life Cycle Assess. 8 (6), 324 330. Joseph, L.M., Poon, D., Shieh, S., 2006. Ingredients of high-rise design: Taipei 101 the World Tallest Building. Structure Magazine 40 45. June 2006. JSSI Manual, 2003. Design and Construction Manual for Passively Controlled Buildings, first ed. Japan Society of Seismic Isolation (JSSI), Tokio, Japan. JSSI Manual, 2005. Design and Construction Manual for Passively Controlled Buildings, second ed. Japan Society of Seismic Isolation (JSSI), Tokio, Japan. JSSI Manual, 2007. Design and Construction Manual for Passively Controlled Buildings, third ed. Japan Society of Seismic Isolation (JSSI), Tokio, Japan. JSSI Standard for Seismic Isolation Devices, 2000, published by JSSI, 2000. JSSI, 2010. JSSI Guideline for Design of Seismically Isolated Buildings using TimeHistory Analysis Method (in Japanese). Jung, H.J., Park, K.S., Spencer, B.F., Lee, W., 2004. Hybrid seismic protection of cablestayed bridges. Earthquake Eng. Struct. Dyn. 33 (7), 795 820. Jurukovski, D., Rakicevic, Z., 1995. Vibration base isolation development and application. In: Proceedings of Tenth European Conference on Earthquake Engineering, Duma, Balkema, Rotterdam, 667 676.

1035

1036

References

Kaczmarczyk, S., 2012. The nonstationary, nonlinear dynamic interactions in slender continua deployed in high-rise vertical transportation systems in the modern built environment. J. Phys. Conf. Ser. 382, 012037. Kaczmarczyk, S., Picton, P., 2013. The prediction of nonlinear responses and active stiffness control of moving slender continua subjected to dynamic loadings in a vertical host structure. Int. J. Acoust. Vib. 18, 39 44. Kaczmarczyk, S., Andrew, J.P., Adams, J.P., 2003. The modelling and prediction of the influence of building vibration on the dynamic response of elevator ropes. Mater. Sci. Forum 440-441, 489 496. Kamae, K., Kawabe, H., Irikura, K., 2004. Strong ground motion prediction for huge subduction earthquakes using a characterized source model and several simulation techniques. In: Proceedings of the 13th WCEE, Vancouver. Kamura, H., Ito, S., Okamoto, H., Homma, A., 1997. Effects of strain rate on performance of beam-to column connection under strong earthquake Part 2 FEM analysis considering of effects of strain rate. In: Summaries of Technical Papers of Annual Meeting Architectural Institute of Japan, Kanto, Japan, Vol.C-1, pp. 427 428. Kanada, M., Kasai, K., Okuma, K., 2002. Innovative passive control scheme: a Japanese 12-story building with stepping columns and viscoelastic dampers. In: Proc. Structural Engineers World Congress (SEWC), Yokohama, Japan. Kanda, J., Tamura, Y., Fujii, K., 1990. Probabilistic perception limits of low-frequency horizontal motions. In: Conference with International Participation, Serviceability of Steel and Composite Structures Proceedings, Pardubice, Czechoslovakia, pp. 61 72. Kaneko, S., Ishikawa, M., 1999. Modelling of a tuned liquid damper with submerged nets. J. Press. Vessel Technol. 121, 334 343. Kani, N., Feng, D., Sera, S., 2010. Performance based design of seismically isolated buildings in japan. In: Workshop: Bridges seismic isolation and large scale modeling, St.Petersburg, Russia, June 29th July 3rd. Kanno, Y., Takewaki, I., 2006. Sequential semidefinite program for maximum robustness design of structures under load uncertainties. J. Optimiz. Theory App. 130 (2), 265 287. Kar, I.N., Miyakura, T., Seto, K., 2000. Bending and torsional vibration control of flexible structure using HN-based robust control law. IEEE Trans. Mech. 8 (3), 545 553. Kareem, A., 1981. Wind-excited response of buildings in higher modes. J. Struct. Eng. ASCE 107 (ST4), 701 706. Kareem, A., Gurley, K., 1996. Damping in structures: its evaluation and treatment of uncertainty. J. Wind Eng. Ind. Aerodyn. 59, 131 157. Kareem, A., Kline, S., 1995. Performance of multiple mass dampers under random loading. J. Struct. Eng. 121 (2), 348 361. Kareem, A., Sun, W.J., 1987. Stochastic response of structure with fluid-containing appendages. J. Sound Vibrat. 119 (3), 389 408. Kareem, A., Kijewski, T., Tamura, Y., 1999. Mitigation of motions of tall buildings with specific examples of recent applications. Wind Struct. 2 (3), 201 251. Kasagi, M., Fujita, M., Takewaki, I., 2016. Automatic generation of smart earthquakeresistant building systems: hybrid system of base-isolation and building-connection. Heliyon 2 (2). Kasai, K., Kibayashi, M., 2004. JSSI manual for building passive control technology part-1 manual contents and design/analysis methods. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1 6, 2004 Paper No. 2989.

References

Kasai, K., Ito, H., Ogura, T., 2008. Passive control design method based on tuning of equivalent stiffness of bilinear oil damper. J. Struct. Constr. Eng. AIJ 73 (630), 1281 1288 (in Japanese). Kasai, K., Ito, H., Ooki, Y., Hikino, T., Kajiwara, K., Motoyui, S., et al., 2010. Full-scale shake table tests of 5-story steel building with various dampers. In: 7th International Conference on Urban Earthquake Engineering (7CUEE) & 5th International Conference on Earthquake Engineering (5ICEE) March 3 5, 2010, Tokyo Institute of Technology, Tokyo, Japan. Kasai, K., Pu, W.C., Wada, A., 2012a. Response of Passively-Controlled Tall Buildings in Tokyo during 2011 Great East Japan Earthquake. 15 World Conference of Earthquake Engineering, Lisboa. Kasai, K., Pu, W., Wada, A., 2012b. Responses of controlled tall buildings in Tokyo subjected to the great east japan earthquake. In: Proceedings of the International Symposium on Engineering Lessons Learned from the 2011 Great East Japan Earthquake, March 1 4, Tokyo, Japan. Kasai, K., Mita, A., Kitamura, H., Matsuda, K., Morgan, T.A., Taylor, A.W., 2013. Performance of seismic protection technologies during the 2011 Tohoku-Oki Earthquake. Earthquake Spectra 29 (1), 256 293. Kasemi, B., Asan, G.A., Muthalif, M.M., Rashid, Rahman, M., 2011. Optimizing dynamic range of magnetorheological fluid dampers: modeling and simulation. In: 4th International Conference on Mechatronics (ICOM), 17-19 May, Kuala Lumpur, Malaysia. Katayama, T., Ito, S., Kamura, H., Ueki, T., Okamoto, H., 2000. Experimental study on hysteretic damper with low yield strength under dynamic loading. In: Proceedings of the 12th World Conference in Earthquake Engineering, Auckland, New Zealand. Kaveh, A., Mohammadi, S., Khadem Hosseini, O., Keyhani, A., Kalatjari, V.R., 2015. Optimum parameters of tuned mass dampers for seismic applications using charged system search. IJST Trans. Civil Eng. 39 (C1), 21 40. Kawabata, I., Takayama, M., Nishikawa, Y., Kimura, Y., Yamazaki, E., Isshiki, Y., 2001. Structural design of high-rise building with base isolation system using elastic sliding bearings and rubber bearings. AIJ J. Technol. Design 12, 99 104. Kelly, J.M., 1986. Aseismic base isolation: review and bibliography. Soil Dyn. Earthquake Eng. 5 (4), 202 216. Kelly, T.E., 2001a. In-Structure Damping and Energy Dissipation Design Guidelines. Holmes Consulting Group Ltd, New Zealand. Kelly, T.E., 2001b. Base Isolation of Structures Design Guidelines. Holmes Consulting Group Ltd, New Zealand. Kelly, J.M., Chalhoub, M.S., 1990. Earthquake simulator testing of a combined sliding bearing and rubber bearing isolation system. In: Report No. UCB/EERS-87-04., Earthquake Engineering Research Center, University of California Berkeley. Kelly, J.M., Skinner, R.I., Heine, A.J., 1972. Mechanism of energy absorption in special devices for use in earthquake resistant structures. Bull. N.Z. Soc. Earthquake Eng. 5 (3), 63 88. Kersting, R.A., Fahnestock, L.A., Lo`pez, W.A., 2015. Seismic Design of Steel bucklingRestrained Braced Frames: A Guide for Practicing Engineers. NEHRP Seismic Design Technical Brief No. 11. Kibayashi, M., Kasai, K., Tsuji, Y., Kikuchi, M., Kimura, Y., Kobayashi, T., et al., 2004. JSSI manual for building passive control technology part-2 criteria for implementation of energy dissipation devices. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1 6, No. 2990.

1037

1038

References

Kidokoro, R., 2004. Self-mass damper (SMD): seismic control system inspired by the pendulum movement of an antique clock. In: The 14 th World Conference on Earthquake Engineering, October 12 17, 2008, Beijing, China. Kijewski-Correa, T., Pirnia, J.D., 2007. Dynamic behavior of tall buildings under wind: insights form full-scale monitoring. Struct. Des. Tall Special Build. 16 (4), 471 486. Kijewski-Correa, T., Kilpatric, J., Kareem, A., Kwon, D.K., Bashor, R., Kochly, M., et al., 2006. Validating wind-induced response of tall buildings: synopsis of the Chicago full scale monitoring program. J. Struct. Eng. 132 (10), 1509. 1509 1523. Kijewski-Correa, T., et al., 2007. Full-scale performance evaluation of tall buildings under winds. In: 12th International Conference on Wind Engineering, International Association for Wind Engineering (IAWE) Atsugi, Kanagawa, Japan, pp. 351 358. Kim, J., Seo, Y., 2003. Seismic design of steel structures with buckling-restrained knee braces. J. Construct. Steel Res. 59, 1477 1497. Kim, J., Choi, H., Min, K.W., 2003. Performance-based design of added viscous dampers using capacity spectrum method. J. Earthquake Eng. 7 (1), 1 24. Kim, Y., You, K., Cho, J., Hong, D., 2006. The vibration performance experiment of tuned liquid damper and tuned liquid column damper. J. Mech. Sci. Technol. 20 (6), 795 805. Kim, H.-J., Christopoulos, C., Tremblay, R., 2004. Experiment characterization of boltstressed non-asbestos organic (NOA) material-to-steel interface. Report No. UT2004-3, Department of Civil Engineering, University of Toronto, Canada. Kim, Y.-M., You, K.-P., Kim, H.-Y., 2008. Wind-induced excitation control of a tall building with tuned mass dampers. Struct. Design Tall Spec. Build. 17 (3), 669 682. Klembczyk, A., 2009. Introduction to shock and vibration isolation and damping systems. In: Proceedings of the Society for Experimental Mechanics, Inc., International Modal Analysis Conference, IMAC XXVII Conference and Exposition on Structural Dynamics. Klembczyk, A., 2014. Utilizing damping devices to improve resiliency of structures. In: Proceedings of the 15th U.S. Japan Workshop on the Improvement of Structural Engineering and Resiliency Applied Technology Council. Klembczyk, A., Breukelman, B., 2000. Structural control of high rise building using a tuned mass with integral hermetically sealed, frictionless hydraulic dampers. In: Proceedings of the 71st Shock and Vibration Symposium. Kneer, E., Maclise, L., 2008. Consideration of building performance in sustainable design: a structural engineer’s role. In: Proceedings of SEAOC Convention, pp. 1 15. Kobori, T., Minai, R., 1960. Analytical study on active seismic response control. Trans. AIJ 66, 253-156 (in Japanese). Kobori, T., Koshika, N., Yamada, K., Ikeda, Y., 1991a. Seismic-response-controlled structure with active mass driver system. part 1: design. Earthquake Eng. Struct. Dyn. 20, 133 149. Kobori, T., Koshika, N., Yamada, K., Ikeda, Y., 1991b. Seismic-response-controlled structure with active mass driver system. Part 2: design. Earthquake Eng. Struct. Dyn. 20, 151 166. Kobori, T., Miura, Y., Fukuzawa, E., Yamada, T., Arita, T., Takenaka, Y., et al., 1992. Development and application of hysteresis steel dampers. In: Proceeding of the 10th World Conference on Earthquake Engineering, Rotterdam, The Netherland. Komuro, T., Nishikawa, Y., Kimura, Y., Isshiki, Y., 2005. Development and realization of Base Isolation System for High-rise Buildings. J. Adv. Concrete Technol. 3 (2), 233 239.

References

Kwan, W.P., Billington, S.L., 2003. Influence of hysteretic behavior on equivalent period and damping of structural systems. J. Struct. Eng. 129 (5), 576 585. Kwok, K.C.S., Samali, B., 1995. Performance of tuned mass dampers under wind loads. Eng. Struct. 17 (9), 655 667. Lago, A., 2011. Seismic Design of Structures with Passive Energy Dissipation Systems. IUSS Report, Pavia, Italy. Lago, A., Sullivan, T.J., Calvi, G.M., 2010. A novel seismic design strategy for structures with complex geometry. J. Earthquake Eng. 14 (S1), 69 105. Lagomarsino, S., 1993. Forecast models for damping and vibration periods of buildings. J. Wind. Eng. Ind. Aerodyn. 48, 221 239. Lagomarsino, S., Pagnini, L.C., 1995. Criteria for modeling and predicting dynamic parameters of buildings. Rep. No. ISC-II, Istituto di Scienza delle Costruzioni, Universita’ of Genova, Facolta’ di Ingegneria. Lagos, R., Boroschek, R., Retamales, R., Lafontaine, M., 2014. Seismic Isolation of the Nunoa Capital Building, the Tallest Base Isolated Residential Building in the Americas. In: Proceedings of the Tenth U.S. National Conference on Earthquake Engineering, Anchorage, Alaska. Lagos, R., Boroschek, R., Retamales, R., Lafontaine, M., Friskel, K., Kasalanati, A., 2017. Seismic Isolation of the Nunoa Capital Building, The Tallest Base Isolated Building in the Americas, 16 WCEE, Santiago, Chile, January 2017. Lai, J.-W., Wang, S., Schoettler, M.J., Mahin, S.A., 2015. Seismic Evaluation and Retrofit of Existing Tall Buildings in California: Case Study of a 35-Story Steel MomentResisting Frame Building in San Francisco. PEER Report No. 2015/14. University of California, Berkeley. Lai, M.L., Chang, K.C., Soong, T.T., Hao, D.S., Yeh, Y.C., 1995. Full-scale viscoelastically damped steel frame. J. Struct. Eng. 121 (10), 1443 1447. Lamb, H., 1932. Hydrodynamics. Cambridge University Press, London. LATBSDC, 2017. An alternative procedure for seismic analysis and design of tall buildings located in the Los Angeles region: a consensus document. Los Angeles Tall Buildings Structural Design Council. Lavan, O., 2012. On the efficiency of viscous dampers in reducing various seismic responses of wall structures. Earthquake Eng. Struct. Dyn. 41 (12), 1673 1692. Lee, D., Ng, M., 2010. Application of tuned liquid dampers for the efficient structural design of slender tall buildings. CTBUH J. 4, 30 36. Lee, D., Taylor, D.P., 2001. Viscous damper development and future trends. Struct. Design Tall Build. 10 (5), 311 320. Lee, G.C., Liang, Z., 2012. Earthquake-Resistant Structures-Design, Assessment and Rehabilitation, Chapter 3: “Design Principles of Seismic Isolation”. InTech. Lee, S.-H., Park, J.-H., Lee, S.-K., Min, K.-W., 2008a. Allocation and slip load of friction dampers for a seismically excited building structure based on storey shear force distribution. Eng. Struct. 30, 930 940. Lee, S.-K., Park, J.-H., Moon, B.-W., Min, K.-W., Lee, S.-H., 2008b. Design of a bracing-friction damper system for seismic retrofitting. Smart Struct. Syst. 4 (5), 685 696. Lee, S.-K., Min, K.-W., Lee, H.-R., 2011. Parameter identification of new bidirectional tuned liquid column and sloshing dampers. J. Sound Vibrat. 330 (7), 1312 1327.

1039

1040

References

Lee, S.-K., Lee, H.-R., Min, K.-W., 2012. Experimental verification on nonlinear dynamic characteristic of a tuned liquid column damper subjected to various excitation amplitudes. Struct. Design Tall Spec. Build. 21 (5), 374 388. Leonard, R.B., 2002. Integrated Buildings: The Systems Basis of Architecture. John Wiley & Sons, New York. Leong, S.L., 2005. Using Shape Memory Alloy as Dampers: Design Methodology. Master Degree Thesis, Massachusetts Institute of Technology, Boston, MA. Lepelletier, T.G., Raichlen, F., 1988. Nonlinear oscillations in rectangular tanks. J. Eng. Mech. 114 (1), 1 23. Leung, A.Y.T., Zhang, H., 2009. Particle swarm optimization of tuned mass dampers. Eng. Struct. 31, 715 728. Levy, R., Lavan, O., 2006. Fully stressed design of passive controllers in framed structures for seismic loadings. Struct. Multidiscipl. Optim. 32 (6), 485 498. Levy, R., Marianchik, E., Rutenberg, A., Segal, F., 2001. A simple approach to the seismic design of friction damped braced medium-rise frames. Eng. Struct. 23, 250 259. Lewandowski, R., Grzymislawska, J., 2009. Dynamic analysis of structures with multiple tuned mass dampers. J. Civil Eng. Manage. 15 (1), 77 86. Li, C., 2000. Performance of multiple tuned mass dampers for attenuating undesirable oscillations of structures under the ground acceleration. Earthquake Eng. Struct. Dyn. 29 (9), 1405 1421. Li, C., 2002. Optimum multiple tuned mass dampers for structures under the ground acceleration based on DDMF and ADMF. Earthquake Eng. Struct. Dyn. 31 (4), 897 919. Li, G., Li, H., 2008. Earthquake-resistant design of RC frame with ‘dual functions’ metallic dampers. In: Proceedings of the 14th World Conference on Earthquake Engineering, Beijing, China. Li, C., Liu, Y., 2002. Further characteristics for multiple tuned mass dampers. J. Struct. Eng. 128 (10), 1362 1365. Li, C., Reinhorn, A.M., 1995. Experimental and analytical investigation of seismic retrofit of structures with supplemental damping: Part II-friction devices. Technical Report NCEER-95-0009, State University of New York at Buffalo, Buffalo, NY. Li, G., Jiang, Y., Zhang, S., Zeng, Y., Li, Qi, 2015. Seismic design or retrofit of buildings with metallic structural fuses by the damage-reduction spectrum. Earthquake Eng. Eng. Vib. 14, 85 96. Li, H.N., Wu, X.X., 2006. Limitations of height-to-width ratio for base-isolated buildings under earthquake. Struct. Design Tall Spec. Build. 15, 277 287. Li, Q.S., Fang, J.Q., Jeary, A.P., Wong, C.K., 1998. Full-scale measurement of wind effects on tall buildings. J. Wind Eng. Ind. Aerodyn. 74 76 (1), 741 750. Li, Q.S., Cao, H., Li, G.Q., Li, S.J., Liu, D.K., 1999. Optimal design of wind-induced vibration control of tall buildings and highrise structures. Wind Struct. 2 (1), 69 83. Li, Q.S., Yang, K., Zhang, N., Wong, C.K., Jeary, A.P., 2002. Field measurements of amplitude dependent damping in a 79-storey tall building and its effects on the structural dynamic response. Struct. Design Tall Build. 11, 129 153. Li, Q.S., Xiao, Y.Q., Wong, C.K., Jeary, A.P., 2003. Field measurements of wind effects on the tallest building in Hong Kong. Struct. Des. Tall Special Build. 12 (1), 67 82. Li, Q.S., Liu, D.K., Tang, J., Zhang, N., Tam, C.M., 2004a. Combinatorial optimal design of number and positions of actuators in actively controlled structures using genetic algorithms. J. Sound Vib. 270 (3 5), 611 624.

References

Li, Q.S., Wu, J.R., Liang, S.G., Xiao, Y.Q., Wong, C.K., 2004b. Full-scale measurements and numerical evaluation of wind induced vibration of a 63-story reinforced concrete tall building. Eng. Struct. 26 (12), 1779 1794. Li, Q.S., Xiao, Y.Q., Wu, J.R., Fu, J.Y., Li, Z.N., 2008. Typhoon effects on super-tall buildings. J. Sound Vib. 313 (3-5), 581 602. Li, T., Zhang, J.Y., Zhang, W.H., Zhu, M.H., 2009. Vortex-induced vibration characteristics of an Elastic Circular Cylinder, World Academy of Science. Eng. Technol. 3. Li, Q.S., Zhi, L.H., Tuan, A.Y., Kao, C.S., Su, S.C., Wu, C.F., 2011. Dynamic behavior of Taipei 101 tower: field measurement and numerical analysis. J. Struct. Eng. 143 155. Available from: https://doi.org/10.1061/(ASCE)ST.1943-541X.0000264. Liang, Z., Lee, G.C., Dargush, G.F., Song, J., 2012. Structural Damping: Applications in Seismic Response Modification. CRC Press, Taylor & Francis Group, Boca Raton, FL. Lim, B., Matsumoto, T., 2006. Characteristics of super high damping visco-elastic damper for earthquake and wind-induced vibration. In: 4th International Conference on Earthquake Engineering, No. 305. Lin, Y.Z., Christenson, R., 2011. Real-time hybrid test validation of a MR damper controlled building with shake table tests. Adv. Struct. Eng. 14 (1), 79 92. Lin, R.C., Liang, Z., Soong, T.T., Zhang, R.H., 1988. An Experimental Study of Seismic Structural Response with Added Viscoelastic Dampers. Report NCEER-88-0018, NCEER, State Univ. of New York at Buffalo. Lin, Y.Y., Tsai, M.H., Hwang, J.S., Chang, K.C., 2003. Direct displacement-based design for building with passive energy dissipation systems. Eng. Struct. 25, 25.37. Lin, P.Y., Chung, L.L., Loh, C.H., 2005. Semiactive control of building structures with semiactive tuned mass damper. Comp. Aided Civil Infrastruct. Eng. 20 (1), 35 51. Lin, Y.Y., Chang, K.C., Chen, C.Y., 2008. Direct displacement-based design for seismic retrofit of existing buildings using non-linear viscous dampers. Bull. Earthquake Eng. 6, 535 552. Lin, C.-C., Wang, J.-F., Lien, C.-H., Chiang, H.-W., Lin, C.-S., 2010. Optimum design and experimental study of multiple tuned mass dampers with limited stroke. Earthquake Eng. Struct. Dyn. 39 (14), 1631 1651. Lin, C.C., Lin, G.L., Lung, H.Y., 2014. Dynamic test of multiple tuned mass dampers for vibration control of high-rise buildings. Tenth U.S. National Conference on Earthquake Engineering Frontiers of Earthquake Engineering, July 21 25, Anchorage, Alaska. Liu, D., Yang, Y.L., Li, Q.S., 2003. Optimum positioning of actuators in tall buildings using genetic algorithm. Comput. Struct. 81 (32), 2823 2827. Liu, T., Zordan, T., Briseghella, B., Zhang, Q., 2014. An improved equivalent linear model of seismic isolation system with bilinear behavior. Eng. Struct. 61, 113 126. Lomiento, G., Bonessio, N., Braga, F., 2010. Design criteria for added dampers and supporting braces. J. Anti-Seismic Syst. Int. Soc. (ASSISI) 1 (1), 55 73. London˜o, J.M., Wagg, D.J., Neild, S.A., 2012. Estimating the overall damping ratio in structures with brace-damper systems. In: Proceedings of the 15th World Conference in Earthquake Engineering, Lisbon, Portugal. Lopez, O.A., Cruz, M., 1996. Number of modes for the seismic design of buildings. Earthquake Eng. Struct. Dyn. 25 (8), 837 856. Lopez-Garcia, D., 2001. A simple method for the design of optimal damper configurations in MDOF structures. Earthquake Spectra 17 (3), 387 398. Love, J., 2016. Innovation: the means to the end. Structure Magazine 50 54.

1041

1042

References

Love, J.S., Tait, M.J., 2010. Non-linear simulation of a tuned liquid damper with damping screens using a modal expansion technique. J. Fluids Struct. 26 (7 8), 1058 1077. Love, J.S., Tait, M.J., 2011. Equivalent linearized mechanical model for tuned liquid dampers of arbitrary tank shape. J. Fluids Eng. 133 (6), 061105. Love, J.S., Tait, M.J., 2012. Equivalent mechanical model for tuned liquid damper of complex tank geometry coupled to a 2D structure. Struct. Control Health Monit. 21, 43 60. Love, J.S., Tait, M.J., 2013. Parametric depth ratio study on tuned liquid dampers: fluid modelling and experimental work. Comput. Fluids 79, 13 26. Love, J., Tait, M., 2015. Multiple tuned liquid dampers for efficient and robust structural control. J. Struct. Eng. 141 (12), 04015045. Lu, X., Chen, J., 2011a. Parameter optimization and structural design of tuned mass damper for shanghai centre tower. Struct. Design Tall Spec. Build. 20 (4), 453 471. Lu, X., Chen, J., 2011b. Mitigation of wind-induced response of Shanghai Center Tower by tuned mass damper. Struct. Design Tall Spec. Build. 20 (4), 435 452. Lu, X., Zhou, Y., Yan, F., 2008. Shaking table test and numerical analysis of RC frames with viscous wall dampers. Struct. Eng. 134 (1), 64 76. Lu, X., Li, P., Guo, X., Shi, W., Liu, J., 2014. Vibration control using ATMD and site measurements on the Shanghai World Financial Center Tower. Struct. Design Tall Spec. Build. 23 (2), 105 123. Lu, X., Zhang, Q., Weng, D., Zhou, Z., Wang, S., Mahin, S.A., et al., 2016. Improving performance of a super tall bulging using a new eddy-current tuned mass damper. Struct. Contr. Health Monit. 24 (3). Luft, R.W., 1979. Optimal tuned mass dampers for buildings. J. Struct. Div. ASCE 2766 2772. Ma, C.F., Zhang, Y.H., Tan, P., Zhou, F.I., 2014. Seismic Response of Base-Isolated HighRise Buildings under Fully Nonstationary Excitation. Shock and Vibration, Hindawi Publishing Corporation. Maddaloni, G., Caterino, N., Occhiuzzi, A., 2016. Shake table investigation of a structure isolated by recycled rubber devices and magnetorheological dampers. Struct. Control Health Monit. Available from: https://doi.org/10.1002/stc.1906. Maebayshi, K., Shiba, K., Mita, A., Inada, Y., 1992. Hybrid Mass damper system for response control of building. In: Proc. of Tenth World Conference on Earthquake Engineering, Rotterdam. Mahmoodi, P., 1969. Structural dampers. J. Struct. Div. 95 (ST8), ASCE, New York. Mahmoodi, P., Robertson, L.E., Yontar, M., Moy, C., Feld, L., 1987. Performance of Viscoelastic Dampers in World Trade Center Towers. Proc. 5th ASCE Structural Congress, Orlando, Florida. Main, J.A., Krenk, S., 2005. Efficiency and tuning of viscous dampers on discrete systems. J. Sound Vibrat. 286 (1 2), 97 122. Marano, G.C., Greco, R., Chiaia, B., 2010. A comparison between different optimization criteria for tuned mass dampers design. J. Sound Vibrat. 329 (23), 4880 4890. Marian, L., Giaralis, A., 2014. Optimal design of a novel tuned mass-damper-inerter (TMDI) passive vibration control configuration for stochastically support-excited structural systems. Probab. Eng. Mech. 38, 156 164.

References

Marioni, A., 2006. The European standard EN 1337 on structural bearings. In: Sixth World Congress on Joints, Bearings and Seismic Systems for Concrete Structures Halifax, 17 21 September, Canada. Marko, J., Thambiratnam, D., Perera, N., 2006. Study of viscoelastic and friction damper configurations in the seismic mitigation of medium-rise structures. J. Mech. Mater. Struct. 1 (6), 1001 1039. Marsh, A.P., Prakash, M., Semercigil, S.E., Turan, O.F., 2010. A shallow-depth sloshing absorber for structural control. J. Fluids Struct. 26 (5), 780 792. Marshall, R., Knapp, G., 2003. Recent advances in technologies, techniques and materials. In: Parker, D., Wood, A. (Eds.), The Tall Building Reference Book. Routledge, New York, pp. 167 180. Martinez-Rodrigo, M., Romero, M.L., 2003. An optimum retrofit strategy for momentresisting frames with nonlinear viscous dampers for seismic applications. Eng. Struct. 25, 913 925. Marukawa, H., Kato, N., Fujii, K., Tamura, Y., 1996. Experimental evaluation of aerodynamic damping of tall buildings. J. Wind Eng. Ind. Aerodyn. 59, 177 190. Mathias, N., Ranaudo, F., Sarkisian, M., 2016. Mechanical amplification of relative movements in damped outriggers for wind and seismic response mitigation. Int. J. High-Rise Build. 5 (1), 51 62. MathWorks, Inc., 2016. MATLAB and Statistics Toolbox Release 2016b. Natick, Massachusetts, United States. Matsuda, K., Kasai, K., Yamagiwa, H., Sato, D., 2012. Responses of Base-Isolated Buildings in Tokyo during the 2011 Great East Japan Earthquake. 15 WCEE, Lisboa. Matsumoto, T., Akita, S., Amasaki, T., 2012. Seismic performance improvement of the existing high-rise building by connecting to its high-rise extension using viscous dampers. In: Proceedings of the 15 th world conference on earthquake engineering Lisbon, Portugal. Matta, E., 2013. Effectiveness of tuned mass dampers against ground motion pulses. Struct. Eng. 139 (2), 188 198. Mayes, R.L., Naeim, F., 2001. The Seismic Design Handbook: Design of Structures with Seismic Isolation. Springer, US, pp. 723 755. Mayes, R.L., Reis, E., 2015. The U.S. Resiliency Council and the Building Rating System. In: Proceedings of second ATC & SEI Conference on Improving the Seismic Performance of Existing Buildings and Other Structures, 10 12 December, San Francisco, California. Mayes, R.L., Brown, A.G., Pietra, D., 2012. Using seismic isolation and energy dissipation to create earthquake-resilient buildings. Bull. N.Z. Soc. Earthquake Eng. 45 (3), 117 122. Mazzolani, F.M., 2004. Competing issues for aluminum alloys in structural engineering. Prog. Struct. Eng. Mater. 6 (4), 185 196. McNamara, R.J., Taylor, D.P., 2003. Fluid viscous dampers for high-rise buildings. Struct. Design Tall Spec. Build. 12 (2), 145 154. McNamara, R.J., Boggs, D.W., Lai, M.-L., Makris, N., Nielsen, E.J., Cermak, J.E. 1997. Tailoring of Damping in Structures: Design Issues. Proceeding Paper. McVerry, G.H., 1980. Structural identification in the frequency domain from earthquake records. Earthquake Eng. Struct. Dyn. 8 (2), 161 180.

1043

1044

References

Melbourne, W.H., 1988. Comfort criteria for wind-induced motion of structures. Struct. Eng. Int. 8 (1), 40 44. Hong Kong and Shanghai. Menicovich, D., Vollen, J., Amitay, M., Letchford, C., DeMauro, E., Rao, A., et al., 2012. A different approach to the aerodynamic performance of tall buildings. CTBUH J. IV, 18 23. Menna, C., Asprone, D., Jalayer, F., Prota, A., Manfredi, G., 2013. Assessment of ecological sustainability of a building subjected to potential seismic events during its lifetime. Int. J. Life Cycle Assess. 18 (2), 504 515. Merrit, S., Uang, C.M., Benzoni, G., 2003. Subassemblage Testing of Star Seismic Buckling Restrained Braces. in report no. TR-2003/04, University of California, San Diego, La Jolla. MGSN, 2005. MGSN 4.19-05: Multifanctional High Rise Buildings and Complexes. Moscow City Building Code. MHRUD, 2012. JG/T209-2012: Dampers for Vibration Energy Dissipation of Buildings. The Ministry of Housing and Urban-Rural Development. MHRUD, 2013. JGJ297-2013: Technical Specifications for Seismic Energy Dissipation of Buildings. The Ministry of Housing and Urban-Rural Development. Midas, 2017. Integrated Solution System for Civil Engineering. Midas Information Technology Co., Ltd. Min, K., Kim, J., Kim, Y., 2014. Design and test of tuned liquid mass dampers for attenuation of the wind responses of a full scale building. Smart Mater. Struct. 23 (4), 45020 45029. Min, K.-W., Song, S.-H., Ahn, S.K., Kim, H.-J., 2004. Characteristics of Tuned Liquid Column Dampers for a 76-story Benchmark Building. CTBUH 2004 Seoul Conference. Korea. 10 13 October. Min, K.-W., Kim, H.-S., Lee, S.-H., Kim, H., Ahn, S.K., 2005. Performance evaluation of tuned liquid column dampers for response control of a 76-storey benchmark building. Eng. Struct. 27 (7), 1101 1112. Miranda, E., 2006. Reflections on the use of elastic or secant stiffness for seismic evaluation and design of structures. In: First European Conference on Earthquake Engineering and Seismology. Geneva, Switzerland. Miranda, E., Aslani, H., 2003. Probabilitst Response Assessment for Building Specific Loss Estimation. Pacific Earthquake Engineering Research Center, Report 2003/03, University of California, Berkeley, CA, USA. Miranda, E., Akkar, S.D., 2006. Generalized interstory drift spectrum. J. Struct. Eng. 132, 840 852. Miranda, J.C., 2005. On tuned mass dampers for reducing the seismic response of structures. Earthquake Eng. Struct. Dyn. 34 (7), 847 865. Mohebbi, M., Joghataie, A., 2012. Designing optimal tuned mass dampers for non-linear frames by distributed genetic algorithms. Struct. Design Tall Spec. Build. 21 (1), 57 76. Mohebbi, M., Rasouli Dabbagh, H., Shakeri, K., 2015. Optimal design of multiple tuned liquid column dampers for seismic vibration control of MDOF structures. periodica polytechnica. Civil Eng. 59 (4), 543 558. Mola, E., Mola, F., Stefopoulos, G., Segato, C., Pozzuoli, C., 2016. The Use of OMA for the Validation of the Design of the Allianz Tower in Milan. Dynamics of Civil Structures, Volume 2, Conference Proceedings of the Society for Experimental Mechanics Series, pp. 313 324.

References

MOLTI, 2000. Notification No. 1446. Standard for specifications and test methods for seismic isolation devices. Ministry of Land, Infrastructure and Transport, Japan. Monjoine, M.J., 1944. Influence of rate of strain and temperature on yield stress of mild steel. J. Appl. Mech. 11, 211 218. Montgomery, M., 2011. Fork Configuration Dampers (FCDs) for Enhanced Dynamic Performance of High-Rise Buildings. PhD Thesis University of Toronto, Canada. Montgomery, M., Christopoulos, C., 2015. Experimental validation of viscoelastic coupling dampers for enhanced dynamic performance of high-rise buildings. J. Struct. Eng. 141 (5), 1 11. 04014145. Moon, K.S., 2005. Dynamic interrelationship between technology and architecture in tall buildings. PhD Thesis, Massachusetts Institute of Technology. Moon, K.S., 2010. Vertically distributed multiple tuned mass dampers in tall buildings: performance analysis and preliminary design. Struct. Design Tall Spec. Build. 19 (3), 347 366. Moon, K.S., 2011. Structural design of double skin facades as damping devices for tall buildings. Proc. Eng. 14, 1351 1358. Moon, Ks, 2013. Dynamic interaction between technology and architectural aesthetics in tall buildings. J. Urban Technol. 20 (2), 3 24. Moon, K.S., 2016. Integrated damping systems for tall buildings: tuned mass damper/double skin facade damping interaction system. Struct. Design Tall Spec. Build. 25 (5), 232 244. Mooneghi, M.A., Kargarmoakhar, R., 2016. Aerodynamic mitigation and shape optimization of buildings: review. J. Build. Eng. 5, 225 235. Moore, R.E., 1966. Interval Analysis, Englewood Cliffs. Prentice-Hall, New Jersey. Morava, B., Haskett, T., Smith, A., 2012. Enhancing the Serviceability Performance of Tall Buildings Using Supplemental Damping Systems, Ingegneria Sismica. Morita, K., Yasunaga, A., Takayama, M., 2014. Experimental study on basic characteristics and fatigue properties of lead damper for seismic isolation system. In: Tenth U.S. National Conference on Earthquake Engineering Frontiers of Earthquake Engineering, July 21 25, Anchorage, Alaska. Morring, F.J., 2013. NASA Technology Stabilizes All Kinds of Structures. Aviation Week & Space Technology. Moutinho, C., 2012. An alternative methodology for designing tuned mass dampers to reduce seismic vibrations in building structures. Earthquake Eng. Struct. Dyn. 41 (14), 2059 2073. MRIT, etc., 2000. The Notification and Commentary on the Structural Calculation Procedures for Building with Seismic Isolation (in Japanese). Mualla, I.H., Belev, B., 2001. Performance of steel frames with a new friction damper device under earthquake excitation. J. Eng. Struct. Mualla, I.H., Belev, B., 2017. Overview of recent projects implementing rotational friction dampers. In: 16th World Conference on Earthquake Engineering, Santiago, Chile. Mulligan, K.J., 2007. Experimental and analytical studies of semi-active and passive structural control of buildings. Ph.D. Thesis, Department of Mechanical Engineering, University of Canterbury, Christchurch, New Zealand. Mulligan, K.J., Chase, J.G., Mander, J.B., Rodgers, G.W., Elliott, R.B., Franco-Anaya, R., et al., 2009. Experimental validation of semi-active resetable actuators in a 1/5 th scale test structure. Earthquake Eng. Struct. Dyn. 38 (4), 517 536.

1045

1046

References

Murudi, M.M., Mane, S.M., 2004. Seismic Effectiveness of Tuned Mass Damper (TMD) for Different Ground Motion Parameters. In: Proceeding of the 13th World Conference on Earthquake Engineering, Vancouver B.C., Canada. Nabid, N., Hajirasouliha, I., Petkovski, M., 2015. Seismic investigation of multi-storey RC frames retrofitted using a friction panel with different slip load distributions. In: The Annual Postgraduate Research Student Conference, Sheffield, UK, 15th April 2015, pp. 43 48. Naeemi, M., Bozorg, M., 2009. Seismic Performance on Knee Braced Frame. World Academy of Science, Engineering and Technology, 26, 976 980. Naeim, F., Kelly, J.M., 1999. Design of Seismic Isolated Structures: From Theory to Practice. John Wiley & Sons, Chichester, UK. Naeim, F., Lew, M., Carpenter, L.D., Youssef, N.F., Rojas, F., Saragoni, G.R., et al., 2011. Performance of tall buildings in Santiago, Chile during the 27 February 2010 offshore Maule, Chile earthquake. Struct. Design Tall Spec. Build. 20 (1), 1 16. Nagarajaiah, S., 2009. Adaptive passive, semiactive, smart tuned mass dampers: identification and control using empirical mode decomposition, Hilbert transform, and shortterm Fourier transform. Struct. Control Health Monit. 16 (7 8), 800 841. Nagarajaiah, S., Varadarajan, N., 2001. Smart variable stiffness control systems. In: Proceedings SPIE 4330, Smart Structures and Materials. Nagarajaiah, S., Varadarajan, N., 2005. Short time Fourier transform algorithm for wind response control of buildings with variable stiffness TMD. Eng. Struct. 27 (3), 431 441. Nagarajaiah, S., Reinhorn, A., Constantinou, M.C., 1991. 3D-BASIS: Nonlinear dynamic analysis of three dimensional base isolated structures. Rep. No. NCEER-91-0005, National Center for Earthquake Engineering Research, State Univ. of New York at Buffalo, NY. Nagase, T., Hisatoku, T., 1992. Tuned-pendulum mass damper installed crystal tower. Struct. Design Tall Spec. Build. 1 (1), 35 56. Nagashima, I., 2001. Optimal displacement feedback control law for active tuned mass damper. Earthquake Eng. Struct. Dyn. 30, 1221 1242. Nagashima, I., Maseki, R., Asami, Y., Hirai, J., Abiru, J., 2001. Performance of hybrid mass damper system applied to a 36-storey high-rise building. Earthquake Eng. Struct. Dyn. 30 (11), 1615 1637. Naguib, M., Salem, F.A., El-Bayoumi, K., 2015. Dynamic analysis of high rise seismically isolated buildings. Am. J. Civil Eng. 3 (2), 43 50. Nakagawa, K., Shimazaki, D., Yoshida, S., Okada, K., 2015. Application of seismic isolation systems in Japanese high-rise buildings. CTBUH J. 2, 36 40. Nakai, M., 2015. Advanced structural technologies for high-rise buildings in Japan. CTBUH J. 2015 (II), 22 29. Nakashima, M., 1993. Extension of on-line computer control method. U.S. Japan Seminar on Development and Future Dimension of Structural Testing Techniques. Nakashima, M., 1995. Strain-hardening Behavior of Shear Panels made of low-yield Steel. I: test. J. Struct. Eng. 121 (12), 1742 1749. Nakashima, M., Masaoka, N., 1999. Real-time on-line test for MDOF systems. Earthquake Eng. Struct. Dyn. 28, 393 420. Nakashima, M., Chusilp, P., 2003. A partial view of Japanese post-Kobe seismic design and construction practices. Earthquake Eng. Eng. Seismol. 4 (1), 3 13.

References

Nakashima, M., Akazawa, T., Tsuji, B., 1995. Strain-hardening Behavior of Shear Panels made of low-yield Steel. II: model. J. Struct. Eng. 121 (12), 1570 1577. Nakashima, M., Kato, H., Takaoka, E., 1999. Development of real-time pseudo dynamic testing. Earthquake Eng. Struct. Dyn. 21, 79 92. Nakata, Y., Hirota, M., Shimizu, T., Iida, T., 2004. JSSI manual for building passive control technology Part-6 design of steel damper. In: 13 th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1 6. Na¨sse´n, J., Holmberg, J., Wadeskog, A., Nyman, M., 2007. Direct and indirect energy use and carbon emissions in the production phase of buildings: an input-output analysis. Energy 9 (32), 1593 1602. NBCC, 1995. National Building Code of Canada 1995: Structural Commentaries (Part 4). National Research Council Canada, Ottawa, Ontario. NBCC, 2010. National Building Code of Canada 2010. National Research Council Canada, Ottawa, Ontario. NCH, 1996. Chilean Standard: Seismic Design of Buildings. NCH, 2013. Chilean Code for Isolated Structures. NEHRP, 2009. Federal Emergency Management Agency FEMA P-750: Recommended Seismic Provisions for New Buildings and Other Structures. Federal Emergency Management Agency ASCE, Washington, DC. NEHRP, 2012. Federal Emergency Management Agency FEMA P-751: 2009 NEHRP Recommended Seismic Provisions: Design Examples. Federal Emergency Management Agency ASCE, Washington, DC. NEHRP, 2015. Federal Emergency Management Agency FEMA P-1050-1, Recommended Seismic Provisions for New Buildings and Other Structures. Federal Emergency Management Agency ASCE, Washington, DC. Nims, D.K., Richter, P.J., Bachman, R.E., 1993. The use of energy dissipating restraint for seismic hazard mitigation. Earthquake Spectra 9 (3), 467 489. NIST, 2012. NIST GCR 12-917-21: Soil-Structure Interaction for Building Structures. U.S. Department of Commerce, National Institute of Standards and Technology, Gaithersburg, MD. Novak, M., 1974. Effect of soil on structural response to wind and earthquake. Erthquake Eng. Struct. Dyn. 3 (1), 79 96. NSCP, 1999. National Structural Code of the Philippines, Association of Structural Engineers of The Philippines Inc. NTC, 2008. Nuove Norme Tecniche per le Costruzioni. Consiglio Superiore dei Lavori Pubblici. NYCBC, 1968. Building Code of the New York City. New York City Department of Buildings. NV, 2009. NV65: Re`gles definissant les effects de la neige et du vent sur les constructions et annexes. CSTB le Futur en Construction. NZS, 2006. NZ3101: Concrete Structures Standard. New Zealand Standard. NZS, 2009. Steel Structure Standard. Standards New Zealand. OBC, 2012. Ontario Building Code. Queen’s Printer for Ontario. OBC, 2006. Ontario Building Code. Ogura, K., Kawabata, I., Hara, T., Komuro, T., Hara, K., Terashima, T., 1997. Sendai MT building. Build. Lett. 1 8.

1047

1048

References

Ok, S.-Y., Song, J., Park, K.-S., 2009. Development of optimal design formula for bi-tuned mass dampers using multi-objective optimization. J. Sound Vibrat. 322 (1 2), 60 77. Okada, K., Yoshida, S., 2014. Structural design of Nakanoshima Festival Tower. Int. J. High-Rise Build. 3 (3), 173 183. Okamoto, S., Kani, N., Higashino, M., Koshika, N., Kimizuka, M., Midorikawa, M., et al., 2002. Recent developments in seismically isolated buildings in Japan. Earthquake Eng. Eng. Vibrat. 1 (2), 213 225. Olson, D.E., Reed, D.A., 2001. A nonlinear numerical model for sloped-bottom tuned liquid dampers. Earthquake Eng. Struct. Dyn. 30, 731 743. ¨ NORM, 1993. O ¨ NORM B 4014-1 -Belastungsannahmen im Bauwesen O Statische Windwirkungen (nicht schwingungsanfa¨llige Bauwerke). OpenSees, 2016. Open System for Earthquake Engineering Simulation: OpenSees. University of California, Berkeley. Available from: http://opensees.berkeley.edu. Ortiz, O., Castells, F., Sonnemann, G., 2009. Sustainability in the construction industry: a review of recent developments based on LCA. Constr. Build. Mater. 23, 28 39. Ou, J., Li, H., 2009. Design approaches for active, semi-active and passive control systems based on analysis of characteristics of active control force. Earthquake Eng. Eng. Vib. 8, 493 506. ¨ zsarıyıldız, S.S., Bozer, A., 2015. Finding optimal parameters of tuned mass dampers. O Struct. Design Tall Spec. Build. 24 (6), 461 475. Padgett, J.E., Tapia, C., 2013. Sustainability of natural hazard risk mitigation: life-cycle analysis of environmental indicators for bridge infrastructure. J. Infrastruct. Syst. 19 (4), 395 408. Palermo, M., Silvestri, S., Landi, L., Gasparini, G., Trombetti, T., 2016. Peak velocities estimation for a direct five-step design procedure of inter-storey viscous dampers. Bull. Earthquake Eng. 14, 599 619. Pall, A., Pall, R.T., 2004. Performance-based design using pall friction dampers- an economical design solution. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1 6, Paper No. 1955. Pall, A.S., Marsh, C., 1982. Response of friction damped braced frames. J. Struct. DivASCE 108 (9), 1313 1323. Pall, A.S., Pall, R.T., 2004. Performance based design using pall friction dampers an economical design solution. In: Proceedings of the 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada. Pall, A.S., Verganelakis, V., Marsh, C., 1987. Friction dampers for seismic control of Concordia library building. In: Proc. 6th Canadian Conference on Earthquake Engineering, Ottawa, Canada. Pan, P., Zamfirescu, D., Nakashima, M., Nakayasu, N., Kashiwa, H., 2005. Base-isolation design practice in Japan: introduction to the post-Kobe approach. J. Earthquake Eng. 9 (1), 147:171. Park, Y.J., Ang, A.H.-S., 1985. Mechanistic seismic damage model for reinforced concrete. J. Struct. Eng. ASCE 111 (4), 722 739. Park, Y.J., Ang, A.H.-S., Wen, Y.K., 1987. Damage-limiting aseismic design of buildings. Earthquake Spectra 3 (1), 1 26. Park, J.-H., Kim, J., Min, K.-W., 2004. Optimal design of added viscoelastic dampers and supporting braces. Earthquake Eng. Struct. Dyn. 33, 465 484.

References

Park, H.G., Kwack, J.H., Jeon, S.W., Kim, W.K., Choi, I.R., 2007. Framed steel plate wall behavior under cyclic lateral loading. J. Struct. Eng. 113 (3), 378 388. Pasquin, C., Leboeuf, N., Pall, R.T., Pall, A., 2004. Friction dampers for seismic rehabilitation of Eaton’s building, Montreal. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1 6, Paper No. 1949. Passoni, C., Belleri, A., Marini, A., Riva, P., 2014. Existing structures connected with dampers: state of the art and future developments. In: Second European Conference on Earthquake Engineering and Seismology, Istanbul, 25 29th August. Pavlou, E., Constantinou, M.C., 2006. Response of nonstructural components in structures with damping systems. J. Struct. Eng. 132, 1108 1117. Pekcan, G., Mander, J.B., Chen, S.S., 1995. The seismic response of a 1:3 scale model R.C. structure with elastomeric spring dampers. Earthquake Spectra 11 (2), 249 267. Pekcan, G., Mander, J.B., Chen, S.S., 1999a. Design and retrofit methodology for building structures with supplemental energy dissipating systems. Report No. MCEER-99-0021, Multidisciplinary Center for Earthquake Engineering Research, State University of New York at Buffalo, Buffalo, NY. Pekcan, G., Mander, J.B., Chen, S.S., 1999b. Fundamental considerations for the design of nonlinear viscous dampers. Earthquake Eng. Struct. Dyn. 28 (11), 1405 1425. PEER/ATC, 2010. PEER/ATC 72-1: Modeling and Acceptance Criteria for Seismic Design and Analysis of Tall Buildings. Applied Technology Council and Pacific Earthquake Engineering Research Center. PEER, 2017. Tall Building Initiative: Guidelines for Performance-Based Seismic Design of Tall Buildings. Pacific Earthquake Engineering Research Center, Report No. 2017/6. Petrini, L., Maggi, C., Priestley, N., Calvi, M., 2008. Experimental verification of viscous damping modeling for inelastic time history analyses. J. Earthquake Eng. 12 (1), 125 145. Pettinga, J.D., Oliver, S., Kelly, T.E., 2013. A design office approach to supplemental damping using fluid viscous dampers. In: Steel Innovations Conference, Christchurch, New Zealand, February 21 22. Peuportier, B.L.P., 2001. Life Cycle Assessment applied to the Comparative Evaluation of Single Family Houses in the French Context. Energy Build. 33 (5), 443 450. Popov, E.P., 1966. Low cycle fatigue of steel beam to column connections. In: Proceeding of International Symposium on the Effects of Repeated Loadings of Materials and Structures, Vol. VI, RILEM 5 Ins. Ing. Mexico. Pozos-Estrada, A., Hong, H.P., Galswothy, J.K., 2010. Serviceability design factors for wind-sensitive structures. Canad. J. Civil Eng. 37, 728 738. PRC National Standard, 2008. Standard for Classification of Seismic Protection of Building Construction (GB 50223-2008), The Ministry of Housing and Urban-Rural Construction of the People’s Republic of China, China Architecture & Building Press, Beijing, China. Preumont, A., 2011. Vibration Control of Active Structures, third ed. Springer. Verlag Berlin Heidelberg, New York. Preumont, A., Seto, K., 2008. Active Control of Structures. John Wiley & Sons. Ltd, Chichester, UK. Priestley, M.N., 1993. Myths and fallacies in earthquake engineering-conflicts between design and reality. Bull. N.Z. Natl. Soc. Earthquake Eng. 26 (3), 329 341.

1049

1050

References

Priestley, M.J.N., 1991. Overview of the PRESSS Research Programme. PCI J. 36 (4), 50 57. Priestley, M.J.N., Grant, D.N., 2005. Viscous damping in seismic design and analysis. J. Earthquake Eng. 9 (2), 229 255. Priestley, M.J.N., Calvi, G.M., Kowalsky, M.J., 2007. Displacement Based Seismic Design of Structures. IUSS Press, Pavia, Italy. Providakis, C.P., 2008. Effect of LRB isolators and supplemented viscous dampers on seismic isolated buildings under near-fault excitations. Eng. Struct. 32, 1187 1198. Providakis, C.P., 2009. Effect of supplemental damping on LRB and FPS seismic isolators under near-fault ground motions. Soil Dyn. Earthquake Eng. 29, 80 90. Pulselli, R.M., Simoncini, E., Pulselli, F.M., Bastianoni, S., 2007. Energy analysis of building manufacturing, maintenance and use: building indices to evaluate housing sustainability. Energy Build 39 (5), 620 628. Purba, R., Bruneau, M., 2009. Finite element investigation and design recommendations for perforated steel plate shear walls. J. Struct. Eng. ASCE 135 (11), 1367 1376. Qian, F., Ding, S., Song, J., Chen, C.-C., 2012. Testing of fluid viscous damper. 15 WCEE, Lisboa, 2012. Qiu, Z.P., 2003. Comparison of static response of structures using convex models and interval analysis method. Int. J. Numer. Meth. Eng. 56 (12), 1735 1753. Quaranta, G., Chakraborty, S., Marano, G.C., 2011. Robust design of tuned liquid column dampers under stochastic ground motion considering fuzzy uncertainties. In: III ECCOMAS Thematic Conference on Computational Methods in Structural Dynamics and Earthquake Engineering. Corfu, Greece, pp. 25 28. Quaranta, G., Mollaioli, F., Monti, G., 2016. Effectiveness of design procedures for linear TMD installed on inelastic structures under pulse-like ground motion. Earthquake Struct. 10 (1), 239 260. Rahimian, A., 2003. Damper-studded diamonds. Engineering News Record, Troy, MI. Rahimian, A., Martinez, R.E., 1998. Chapultepec Tower, Proceedings National Steel Construction Conference, AISC. Rakicevic, Z., Bogdanovic, A., Jurukovski, D., 2009. Shaking table effectiveness testing of GERB PDD (prestressed damping device) control system. IZIIS Report 2009-40: Skopje, R. Macedonia. Rakicevic, Z.T., Bogdanovic, A., Jurukovski, D., Nawrotzki, P., 2012. Effectiveness of tune mass damper in the reduction of the seismic response of the structure. Bull Earthquake Eng. 10 (3), 1049 1073. Ramadhan, G., 2014. Improving the seismic performance of diagrid steel structures using friction mass damper. Master of Science Thesis, Oregon State University. Ramirez, O.M., Constantinou, M.C., Kircher, C.A., Whittaker, A.S., Johnson, M.W., Gomez, J.D., et al., 2001. Development and Evaluation of Simplified Procedures for Analysis and Design of Buildings with Passive Energy Dissipation Systems, MCEER Report 00-0010 Revision 1, MCEER, University at Buffalo, Buffalo, NY. Rana, R., Soong, T.T., 1998. Parametric study and simplified design of tuned mass dampers. Eng. Struct. 20 (3), 193 204. Randall, S.E., Halsted, D.M., Taylor, D.L., 1981. Optimum vibration absorbers for linear damped systems. J. Mech. Design ASME 103, 908 913. Rao, R.S., Gergely, P., White, R.N., 1995. Retrofit of Non-Ductile Reinforced Concrete Frames Using Friction Dampers. Technical Report NCEER-95-0020.

References

Rasouli, S.K., Yahyai, M., 2002. Control of response of structures with passive and active tuned mass dampers. Struct. Design Tall Build. 11 (1), 1 14. Rawat, A., Ummer, N., Matsagarf, V., 2015. Performance of elliptically rolling rods for base isolation of buildings under near-fault earthquakes. In: Proceedings of the 14th World Conference on Seismic Isolation, Energy Dissipation and Active Vibration Control of Structures, September 9-11 2015 San Diego, California, USA. REDI, 2013. Redi Rating System. Resilience-based Earthquake Design Initiative for the Next Generation of Buildings, Version 1.0. Reinhorn, A.M., Soong, T.T., Riley, M., Lin, R.C., Aizawa, A., Higashino, M., 1993. Fullscale implementation of active control. ii: installation and performance. J. Struct. Eng. 119 (6), 1935 1960. Reinhorn, A.M., Nagarajaiah, S., Constantinou, M.C., Tsopelas, P., Li, R., 1994. 3DBASIS-TABS (Version 2.0): computer program for nonlinear dynamic analysis of three-dimensional base isolated structures. Rep. No. NCEER-94-0018, National Center for Earthquake Engineering Research, State Univ. of New York at Buffalo, NY. Reis, E., Mayes, R.L., 2015. The U.S. Resiliency Council’s (USRC) Goals, Objectives, and Founding Principles. Proceedings of second ATC & SEI Conference on Improving the Seismic Performance of Existing Buildings and Other Structures. December 10 12, San Francisco, California. Restrepo-Posada, J.I., 1993. Seismic Behaviour of Connections Between Precast Concrete Elements. Research Report No. 93-3, Department of Civil Engineering, University of Canterbury, New Zealand. Restrepo-Posada, J.I., Dodd, L.L., Park, R., Cooke, N., 1994. Variable affecting cyclic behaviour of reinforcing steel. ASCE J. Struct. Eng. 120 (11), 3178 3196. Rezaeian S., Bozorgnia Y., Idriss I.M., Campbell K.W., Abrahamson N.A., Silva W.J. 2012. Spectral damping scale factors for shallow crustal earthquakes in active tectonic regions, PEER Report No. 2012/01, Pacific Earthquake Engineering Research Center, University of California, Berkeley, CA. Ribakov, Y., Gluck, J., Reinhorn, A.M., 2001. Active viscous damping system for control of MDOF structures. Earthquake Eng. Struct. Dyn. 30, 195 212. Rice, R.C., Jackson, J.L., Bakuckas, J., Thompson, S., 2003. Metallic Materials Properties Development and Standardization. U.S. Department of Transportation, Washington, DC. Ricles, J.M., Sause, R., Garlokc, M., Zhao, C., 2001. Post-tensioned seismic-resistant connections for steel frames. ASCE J. Struct. Eng. 127 (2), 113 121. Ritter, A., 2007. Smart Materials in Architecture. Interior Architecture and Design. Birkhauser, Basel. Robinson, W.H., Greenbank, L.R., 1975. Properties of an extrusion energy absorber. Bull. N.Z. Soc. Earthquake Eng. 8 (3), 187 191. Robinson, W.H., Greenbank, L.R., 1976. Extrusion energy absorber suitable for the protection of structures during an earthquake. Earthquake Eng. Struct. Dyn. 4 (3), 251 259. Robinson, W.H., Cousins, W.J., 1987. Recent development in lead dampers for base isolation. In: Pacific Conference on Earthquake Engineering, Vol. 2, New Zealand. Rodgers, G.W., Chase, J.G., Mander, J.B., Leach, N.C., Denmead, C.S., 2006. Experimental development, tradeoff analysis and design implementation of high force-to-volume extrusion damper technology. Bull. N.Z. Natl. Soc. Earthquake Eng. 40 (2), 35 48.

1051

1052

References

Rodgers, G.W., Mander, J.B., Chase, J.G., Mulligan, K.J., Deam, B.L., Carr, A., 2007. Reshaping hysteretic behaviour—spectral analysis and design equations for semi-active structures. Earthquake Eng. Struct. Dyn. 36 (1), 77 100. Roffel, A.J., 2012. Condition Assessment of In-Service Pendulum Tuned Mass Dampers. PhD Thesis. University of Waterloo. Roffel, A.J., Lourenco, R., Narasimhan, S., Yarusevych, S., 2011. Adaptive compensation for detuning in pendulum tuned mass dampers. J. Struct. Eng. 137 (2), 242 251. Ross, D., 2004. HVAC Design Guide for Tall Commercial Buildings. American Society of Heating, Refrigerating and Air-Conditioning Engineers, Inc. USA. Roussis, P.C., Constantinou, M.C., 2006. Uplift-restraining friction pendulum seismic isolation system. Earthquake Eng. Struct. Dyn. 35 (5), 577 593. Rozas, L., Boroschek, R.L., Tamburrino, A., Roja, M., 2016. A bidirectional tuned liquid column damper for reducing the seismic response of buildings. Struct. Control Health Monit. 23, 621 640. Ruiz, R.O., Lopez-Garcia, D., Taflanidis, A.A., 2016. Modeling and experimental validation of a new type of tuned liquid damper. Acta Mech. 227 (11), 1 20. SAC, 1996. Analytical and Field Investigations of Buildings Affected by the Northridge Earthquake, Report No. SAC-95-04, prepared by SAC Joint Venture, a partnership of Structural Engineers Association of California, Applied Technology Council, and California Universities for Research in Earthquake Engineering. Sadek, F., Mohraz, B., Taylor, A.W., Chung, R.M., 1996. Passive Energy Dissipating Devices for Seismic Applications. Report NISTIR 5923. Building and Fire Research Laboratory, National Institute of Standards and Technology, Gaithersburg, Maryland. Sadek, F., Mohraz, B., Taylor, A.W., Chung, R.M., 1997. A method of estimating the parameters of tuned mass dampers for seismic applications. Earthquake Eng. Struct. Dyn. 26 (6), 617 635. Sadek, F., Mohraz, B., Lew, H.S., 1998. Single- and multiple-tuned liquid column dampers for seismic applications. Earthquake Eng Struc. Dyn. 27, 439 463. Safak, E., Celebi, M., 1991. Seismic response of Transamerica building. II: system identification. J. Struct. Eng. 117 (8), 2405 2425. Saha, S., Debbarma, R., 2017. An experimental study on response control of structures using multiple tuned liquid dampers under dynamic loading. Int. J. Adv. Struct. Eng. 9, 27 35. Sakai, F., Takeada, S., Tamaki, T., 1989. Tuned liquid column damper-new type devices for suppression of building vibration. In: Proceedings of the International Conference on High-Rise Buildings, pp. 926-941, China. Samanta, A., Banerji, P., 2012. Earthquake vibration control using sloshing liquid dampers in building structures. J. Earthquake Tsunami 6 (1), 250002-(1:25). SAMCO, 2006. F08b guideline for structural health monitoring. In: Ru¨cker, W., Hille, F., Rohrmann, R. (Eds.), SAMCO final report. BAM Federal Institute of Materials Research and Testing, Berlin, Germany (2006). Samui, P., Chakraborty, S., Kim, D., 2016. Modeling and Simulation Techniques in Structural Engineering. IGI Global. Sarlis, A.A.S., Constantinou, M.C., 2010. Modeling Triple Friction Pendulum Isolators in Program SAP2000. Sasaki, A., Suganuma, S., Suda, K., Tamura, Y., 1998. Full-scale database on dynamic properties of buildings—Frequency and amplitude dependencies of buildings. In: Proc.,

References

Annual Meeting of the Architectural Institute of Japan (AIJ), Fukuoka, AIJ Japan, B-2, pp. 379 380. Satake, N., Suda, K., Arakawa, T., Sasaki, A., Tamura, Y., 2003. Damping evaluation using full- scale data of buildings in Japan. J. Struct. Eng. 470 477. Sato, D., Ooki Y., Morikawa, H., Yamada, S., Sakata, H., Yamanaka, H., et al., 2008. Evaluation of seismically isolated tall building based on long-term monitoring. In: The 14 th World Conference on Earthquake Engineering, October 12 17, Beijing, China. Scheuer, C., Keoleian, G., Reppe, P., 2003. Life cycle energy and environmental performance of a new university building: modeling challenges and design implications. Energy Build. 35, 1049 1064. Scholl, R.E., 1975. Effects prediction guidelines for structures subjected to ground motion. Rep. No. JAB-99-115, URS/Blume Engineers, San Francisco, CA. SEAOC, 1995. Vision 2000: Performance Based Seismic Engineering of Buildings, San Francisco. SEAONC, 2015. Earthquake Performance Rating System User’s Guide. The Building Ratings Committee. Seto, K., Yamashita, S., Ohkuma, M., Nagamatu, A., 1987. Method of estimating equivalent mass of multi-degree-of-freedom system. JSME Int. J. Ser. C Mech. Syst. Mach. Elements Manuf. 30, 1636 1644. Seto, K., Kondo, S., Ezure, K., 1995. Vibration control method for flexible structures with distributed parameters using a hybrid dynamic absorber. ASME 1995 Design Engineering Technical Conf., DE-Vol. 83-3. Seto, K., Ren, M., Doi, F., 1998. Feedback vibration control of a flexible plate at audio frequencies by using a physical state space approach. J. Acoust. Soc. Am. 103, 924 934. SFBC, 2001. San Francisco Building Code. California Building Standards Commission. SFBC, 2013. San Francisco Building Code. California Building Standards Commission, Illinois, USA. UBC, 1979. Uniform Building Code, International Conference of Building Official, Whittier, California, USA. SFDBI, 2007. Requirements and guidelines for the seismic design and review of new tall buildings using non-prescriptive seismic- design procedures. Admin. Bull. 83. San Francisco Department of Building Inspection, San Francisco, California. Shakal, F.A., Huang, M.J., Ventura, C.E., 1988. The California Strong Motion Instrumentation Program: Objectives, Status and Recent Data. Proceedings of Ninth World Conference on Earthquake Engineering, Tokyo, Japan. Sharabash, A.M., Andrawes, B.O., 2009. Application of shape memory alloy dampers in the seismic control of cable-stayed bridges. Eng. Struct. 31, 607 619. Shi, W., Shan, J., Lu, X., 2012. Modal identification of Shanghai World Financial Center both from free and ambient vibration response. Eng. Struct. 36, 14 26. Shih, M.H., Sung, W.P., 2005. A model for hysteretic behaviour of rhombic low yield strength steel added damping and stiffness. Comp. Struct. 83, 895 908. Shih, M. H., Sung, W.-P., Go, C.-G., 2004. Investigation of newly developed added damping and stiffness device with low yield strength steel. J. Zhejiang. Univ. Sci. 5 (3), 326 334. Shimizu, T., Hayama, S., 1987. Nonlinear response of sloshing based on the shallow water wave theory. JSME Int. J. 30, 806 813.

1053

1054

References

Shin, H., 2010. Life-Cycle Cost Based Optimal Seismic Design of Structures with Energy Dissipation Devices. Ph.D. Thesis, Virginia Polytechnic Institute and State University. Shin, J., Lee, K., Jeong, S.H., Lee, H.S., 2012. Experimental study on buckling-restrained knee brace with steel channel sections. In: Proceedings of the 15th World Conference on Earthquake Engineering, on CD. Shinozaki, Y., Hosozawa, O., Komuro, T., 2004. Structural design of base-isolation system for tall building in Japan. In: CTBUH Seoul Conference. Seoul, Korea, 10 13 October. Shioya, K., Kanda, J., 1993. Human perception thresholds of horizontal motion. In: Proceedings of International Colloquium on Structural Serviceability of Buildings, IABSE Reports, vol. 69, Goteborg, Sweden, pp. 45 52. Shum, K.M., 2009. Closed form optimal solution of a tuned liquid column dampers for suppressing harmonic vibration of structures. Eng. Struct. 31 (1), 84 92. Silvestri, S., Gasparini, G., Trombetti, T., 2010. A five-step procedure for the dimensioning of viscous dampers to be inserted in building structures. J. Earthquake Eng. 14 (3), 417 447. Sinclair, M., 2006. Viscous damping devices: design and implementation issues. In: Proceedings of the 100th anniversary conference commemorating the 1906, San Francisco Earthquake, 18 22 April, San Francisco. Singh, M.P., Singh, S., Moreschi, L.M., 2002. Tuned mass dampers for response control of torsional buildings. Earthquake Eng. Struct. Dyn. 31, 749 769. Skinner, R.I., Kelly, J.M., Heine, A.J., 1974. Energy absorption devices for earthquake resistant structures. In: Proceeding 5 th World on Earthquake Engineering. Rome, Vol. 2, pp. 2924 2933. Skinner, R.I., Beck, J.L., Bycroft, G.N., 1975a. A practical system for isolating structures from earthquake attack,”. Earthquake Eng. Struct. Dyn. 3 (3), 297 309. Skinner, R.I., Kelly, J.M., Heine, A.J., 1975b. Hysteretic dampers for earthquake-resistant structures,”. Earthquake Eng. Struct. Dyn. 3 (3), 287 296. Skinner, R.I., Tyler, R.G., Heine, A.J., Robinson, W.J., 1980. Hysteretic dampers for the protection of structures from earthquakes. Bull. N.Z. Soc. Earthquake Eng. 13 (1), 22 36. Skinner, R.I., Robinson, W.H., McVerry, G.H., 1993. An Introduction to Seismic Isolation. John Wiley & Sons Ltd, New York, NY. Smith, M.C., 2002. Synthesis of mechanical networks: the inerter. IEEE Trans. Automat. Control 47 (10), 1648 1662. Smith, R.J., 2011. Deflection limits in tall buildings are they useful? Structures Congress 2011, Las Vegas, Nevada. Smith, R.J., 2016. The damped outrigger design and implementation. Int. J. High-Rise Build. 5 (1), 63 71. Smith, R.J., Willfort, M.R., 2007. The damped outrigger concept for tall building. Struct. Design Tall Spec. Build. 16, 501 517. Smith, R.J., Merello, R., Willford, M.R., 2010. Intrinsic and supplementary damping in tall buildings. Proc. Inst. Civil Eng. Struct. Build. 163 (2), 111 118. Soong, T.T., Constantinou, M.C., 1994. Passive and Active Structural Vibration Control in Civil Engineering. Springer-Verlag, New York. Soong, T.T., Dargush, G.F., 1997. Passive Energy Dissipation Systems in Structural Engineering. John Wiley & Sons, Ltd, UK.

References

Soong, T.T., Spencer, B.F., 2000. Active, semi-active and hybrid control of structures. In: Proceedings of the 12th World Conference in Earthquake Engineering, Auckland, New Zealand. Soong, T.T., Spencer, B.F., 2002. Supplemental energy dissipation: state-of-the-art and state-of-the-practice. Eng. Struct. 24, 243 259. Soong, T.T., Reinhorn, A.M., Wang, R.C., Lin, R.C., Riley, M., 1992. Full-scale implementation of active structural control. In: Earthquake Engineering, Tenth World Conference, Balkema, Rotterdam, pp. 2131 2136. Sorace, S., Terenzi, G., 2009. Fluid viscous damper-based seismic retrofit strategies of steel structures: general concepts and design applications. Adv. Steel Const. 5 (3), 325 342. Soto, M.G., Adeli, H., 2013. Tuned mass dampers. Arch. Comput. Methods Eng. 20 (4), 419 431. Spence, S.M.J., Kareem, A., 2014. Tall buildings and damping: a concept-based datadriven model. J. Struct. Eng. 140 (5). Spencer, B.F., 2002. Smart Damping Technologies for Dynamic Hazard Mitigation. Nist Special Publication, pp. 75 82. Spencer, B.F., Sain, M.K., 1997. Controlling buildings: a new frontier in feedback. IEEE Cont. Syst. Mag. Emerg. Technol. 17 (6), 19 35. Spencer, B.F., Dyke, S.J., Sain, M.K., Carlson, J.D., 1997. Phenomenological model of a magnetorheological damper. J. Eng. Mech. 123 (3), 230 238. Stanton, J.F., Nakaki, S.D., 2002. Design Guidelines for Precast Concrete Structural Systems. PRESS Report No. 01/03-09. Staudacher, E., Habacher, C., Siegenthaler, R., 1970. Erdebensicherung in Baum. Neue Zu¨rcher Zeitung, Technikbeilage, Zu¨rich, Switzerland. Steinbuch, R., 2011. Bionic optimisation of the earthquake resistance of high buildings by tuned mass dampers. J. Bionic Eng. 8, 335 344. Stiemer, S.F., Godden, W.G., Kelly, J.M., 1981. Experimental Behaviour of Spatial Piping System with Steel Energy Absorbers Subjected to a Simulated Differential Seismic Input. Report No. UCB/EERC-81/09, EERC, Univ. of California at Berkeley. Strakosch, G.R., Caporale, R.S., 2010. The Vertical Transportation Handbook. John Wiley & Sons, Hoboken, New Jersey. Su, J.Z., Xia, Y., Chen, L., Zhao, Z., Zhang, Q.L., Xu, Y.L., et al., 2013. Long-term structural performance monitoring system for the shanghai tower. J. Civil Struct. Health Monit. 3 (1), 49 61. Su, J.Z., Xia, Y., Xu, Y.L., Zhao, X., Zhang, Q.L., 2014. Settlement monitoring of a supertall building using the kalman filtering technique and forward construction stage analysis. Adv. Struct. Eng. 17 (6), 881 893. Sueoka, T., Torii, S., Tsuneki, Y., 2004. The application of response control design using middle-story isolation system to high-rise building. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1 6, Paper No. 3457. Suhardjo, J., Spencer Jr., B.F., Sain, M.K., 1990. Feedback-feedforward control of structures under seismic excitation. Struct. Safety 8, 69 89. Sullivan, T.J., Lago, A., 2012. Towards a simplified direct DBD procedure for the Seismic design of moment resisting frames with viscous damper. Eng. Struct. 35, 140 148. Sullivan, T.J., Priestley, M.J.N., Calvi, G.M., 2008. Estimating the higher-mode response of ductile structures. J. Earthquake Eng. 12 (3), 456 472.

1055

1056

References

Sun, L.M., Fujino, Y., Pacheco, B.M., Chaiseri, P., 1992. Modelling of tuned liquid damper. J. Wind Eng. Ind. Aerodyn. 43 (3), 1883 1894. Sun, L.M., Fujino, Y., Chaiseri, P., Pacheco, B.M., 1995. The properties of a tuned liquid damper using TMD analogy. Earthquake Eng. Struct. Dyn. 17 (9), 609 612. Symans, M.D., Constantinou, M.C., 1997a. Seismic testing of a building structure with a semi-active fluid damper control system. Earthquake Eng. Struct. Dyn. 26, 759 777. Symans, M.D., Constantinou, M.C., 1997b. Experimental testing and analytical modeling of semi-active fluid dampers for seismic protection. J. Intell. Mater. Syst. Struct. 8 (8), 644 657. Symans, M.D., Charney, F.A., Whittaker, A.S., Constantinou, M.C., Kircher, C.A., Johnson, M.W., et al., 2008. Energy Dissipation Systems for Seismic Applications: Current Practice and Recent Developments. J. Struct. Eng. 134:1 (3), 3 21. Taflanidis, A.A., Angelides, D.C., Manos, G.C., 2005. Optimal design and performance of liquid column mass dampers for rotational vibration control of structures under white noise excitation. Eng. Struct. 27, 524 534. Taghavi, S., Miranda, E., 2003. Response Assessment of Nonstructural Elements. PEER Report 2003/05, Pacific Earthquake Engineering Research Center, Berkeley. Tait, M., EI Damatty, A.A., Isyumov, N., 2004. Testing of tuned liquid damper with screens and development of equivalent TMD model. Wind Struct. 7 (4), 215 234. Tait, M.J., 2004. The performance of 1-D and 2-D tuned liquid dampers. Ph.D. thesis, The University of Western Ontario. Tait, M.J., 2008. Modelling and preliminary design of a structure-TLD system. Eng. Struct. 30, 2644 2655. Tait, M.J., Deng, X., 2010. The performance of structure-tuned liquid damper systems with different tank geometries. Struct. Control Health Monit. 17, 254 277. Tait, M.J., Isyumov, N., El-Damatty, A.A., 2004. The efficiency and robustness of a unidirectional tuned liquid damper and modelling with an equivalent TMD. Wind Struct. 7 (4), 235 250. Tait, M.J., El-Damatty, A.A., Isyumov, N., Sidique, M., 2005. Numerical flow models to simulate tuned liquid dampers (TLD) with slat screens. J. Fluids Struct. 20 (8), 1007 1023. Tait, M.J., Isyumov, M., El Damatty, A.A., 2008. Performance of tuned liquid dampers. J. Eng. Mech. 134 (5), 417 427. Takeuchi, T., Hajjar, J., Matsui, R., Nishimoto, K., Aiken, I., 2010. Local buckling restraint condition for core plates in buckling restrained braces. J. Construct. Steel Res. 66, 139 149. Takewaki, I., 1997. Optimal damper placement for minimum transfer functions. Earthquake Eng. Struct. Dyn. 26 (11), 1113 1124. Takewaki, I., 2005. Uncertain-parameter sensitivity of earthquake input energy to baseisolated structure. Struct. Eng. Mech. 20 (3), 347 362. Takewaki, I., 2006. Critical Excitation Methods in Earthquake Engineering. Elsevier Science, Amsterdam. Takewaki, I., 2008a. Robustness of base-isolated high-rise buildings under code-specified ground motions. Struct. Design Tall Spec. Build. 17 (2), 257 271. Takewaki, I., 2008b. Critical excitation methods for important structures. Invited as a Semi-Plenary Speaker, EURODYN 2008, July 7 9, Southampton, England. Takewaki, I., 2009. Building Control with Passive Dampers: Optimal Performance-Based Design for Earthquakes. John Wiley & Sons (Asia), Singapore.

References

Takewaki, I., 2013. Toward greater building earthquake resilience using concept of critical excitation: a review. Sust. Cities Soc. 9, 39 53. Takewaki, I., 2015. Beyond uncertainties in earthquake structural engineering. Front. Built Environ. 1 (1), 1 6. Takewaki, I., Fujita, K., 2009. Earthquake input energy to tall and base-isolated buildings in time and frequency dual domains. Struct. Design Tall Spec. Build. 18 (6), 589 606. Takewaki, I., Fujita, K., 2014. Robust control of building structures under uncertain conditions. In: Beer, M., Patelli, E., Kougioumtzoglou, I., Siu-Kui, I. (Eds.), Encyclopedia of Earthquake Engineering. Springer-Verlag, Berlin, Heidelberg. Takewaki, I., Moustafa, A., Fujita, K., 2013a. Earthquake Resilience of High-Rise Buildings: Case Study of the 2011 Tohoku (Japan) Earthquake, Improving the Earthquake Resilience of Buildings. Springer, London. Takewaki, I., Moustafa, A., Fujita, K., 2013b. Improving the Earthquake Resilience of Buildings: The Worst Case Approach. Springer, London. Tamboli, A., 2005. Manhattan’s mixed construction skyscrapers with tuned liquid and mass. In: CTBUH 2005 7 th World Congress, New York. Tamura, Y., 2003. Design issues for tall buildings from accelerations to damping Tribute to Hatsuo Ishizaki and Vinod Modi. In: Proc. 11th International Conference on Wind Engineering, INV.W2,Lubbock, Texas, USA, pp. 81 114. Tamura, Y., 2005. Lecture 10: Damping in Buildings [Lecture Slides] Retrieved 02-02-2016, 2016, from http://www.wind.arch.t-kougei.ac.jp/info_center/ITcontent/tamura/10.pdf. Tamura, Y., 2006. Amplitude dependency of damping in buildings and estimation techniques. In: Proc. 12th Australasian Wind Engineering Society Workshop, Queenstown, New Zealand. Tamura, Y., 2009. Wind and tall buildings. In: Proceedings 5th European and African Conference on Wind Engineering, Florence, Italy. Tamura, Y., 2012. Amplitude dependency of damping inn buildings and critical tip drift ratio. Int. J. High-Rise Build. 1 (1), 1 13. Tamura, Y., Suganuma, S., 1996. Evaluation of amplitude-dependent damping and natural frequency of buildings during strong wings. J. Wind Eng. Ind. Aerodyn. 59, 115 130. Tamura, Y., Suda, K., Sasaki, A., 2000. Damping in buildings for wind resistant design. In: Proc. International Symposium on Wind and Structures for the 21st Century, Cheju, Korea, pp. 115 130. Tamura, Y., Kareem, A., 2013. Advanced Structural Wind Engineering. Springer, New York. Tamura, Y., Fujii, K., Ohtsuki, T., Wakahara, T., Kohsaka, R., 1995. Effectiveness of tuned liquid damper under wind excitation. Eng. Struct. 1995 (17), 609 621. Tan, P., Dyke, S.J., Richardson, A., Abdullah, M., 2005. Integrated device placement and control design in civil structures using genetic algorithms. J. Struct. Eng. 131, 1489 1496. Tanaka, K., Sasaki, Y., 2000. Hysteretic performance of shear panel dampers of ultra lowyield-strength steel for seismic response control of buildings. In: Proceedings of the 12th World Conference in Earthquake Engineering, Auckland, New Zealand. Tanaka, T., Yamamoto, M., Katayama, T., Nakahira, K., Yamane, K., Shimano, Y., et al., 2003. Recent applications of structural control systems to high-rise buildings. Earthquake Eng. Eng. Seismol. 4 (1), 75 93.

1057

1058

References

Tanaka, Y., Kawaguchi, S., Sukagawa, M., Masaki, N., Sera, S., Washiyama, Y., et al., 2004. JSSI manual for building passive control technology part-4 performance and quality control of viscous dampers. In: 13 th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1 6. Tanaka, H., Tamura, Y., Ohtake, K., Nakai, M., Hayano, Y., Koshika, N., 2009. Study on wind force characteristics for super high-rise building with new shape, (Part 1) Comparison of wind force characteristics of various building shapes. In: Summaries of Technical Papers, Annual Meeting, Architectural Institute of Japan, B-1, Structures I, (to be published) (in Japanese). Tanaka, H., Tamura, Y., Ohtake, K., Nakai, M., Kim, Y.C., Bandi, E.K., 2013. Aerodynamic and flow characteristics of tall buildings with various unconventional configurations. Int. J. High-Rise Build. 2 (3), 213 228. Tanzo, W., Yamada, Y., Iemura, H., 1992. Substructured Computer-Actuator Hybrid Loading Tests for Inelastic Earthquake Response of Structures, Research Report No. 92-ST-01, Department of Civil Engineering, Kyoto University. Taylor, D., 1999. Buildings: design for damping. In: Proceedings of the Boston Society of Civil Engineers, BSCES Fall 1999 Lecture Series, “Dynamics for Structures”. Taylor, D., 2002. Taylor Devices Hermetic Dampers: Description, Applications, and Design. Taylor, A.W., 2012. Response control systems in the United States and lessons learned from the Tohoku earthquake. In: Proceedings of the International Symposium on Engineering Lessons Learned from the 2011 Great East Japan Earthquake, March 1 4, Tokyo, Japan. Taylor, D., Constantinou, M.C., 1995. Testing procedures for high output fluid viscous dampers used in building and bridge structures to dissipate seismic energy. Shock Vib. 2 (5), 373 381. Taylor, D., Katz, I., 2007. Seismic Protection with Fluid Viscous Dampers for the Torre Mayor, A 57- Story Office Tower in Mexico City, Mexico. In: Proceedings of the International Symposium under the Auspices of the European Association for Earthquake Engineering, Bucharest, Romania. Taylor, D.P., 1999. Buildings: Design for Damping. Taylor Device Inc. Taylor, D.P., 2010. Smart buildings and viscous dampers a design engineer’s perspective. Struct. Design Tall Spec. Build. 19 (4), 369 372. Taylor, D.P., 2011. Seismic Testing of Full Scale Seismic Testing of Full Scale Fluid Viscous Dampers Fluid Viscous Dampers. Taylor Device Inc. NEES-MCEER Annual Meeting, Hyatt Regency Hotel Buffalo, NY, June 9 11. Taylor, D.P., Lee, D.A., 1987. Precise Positioning Shock Isolators. 60th Shock and Vibration Symposium, Huntsville, Alaska. Taylor, D.P., Constantinou, M.C., 1996. Fluid Dampers for Applications of Seismic Energy Dissipation and Seismic Isolation. Publication of Taylor Devices, Inc. Taylor, D.P., Constantinou, M.C., 2000. Fluid Dampers for Applications of Seismic Energy Dissipation and Seismic Isolation. Taylor Device Inc. Taylor, A.W., Lin, A.N., Martin, J.W., 1992. Performance of elastomers in isolation bearings: a literature review. Earthquake Spectra 8, 279 303. Taylor, A.W., Shenton, H.W., Chung, R.M., 1995. Standards for the testing and evaluation of seismic isolation systems. In: Proceedings of the ASME/JSME Pressure Vessels and Piping Conference, Honolulu, Hawaii, 319, 39-43.

References

Taylor, D.P., Constantinou, M.C., Symans, M.D., Tsopelas, P., 2000. Fluid Viscous Dampers in Applications of Seismic Energy Dissipation and Seismic Isolation. Taylor Devices, Technical Paper. TDY, 2007. Turkish Earthquake Code: Specification for buildings to be built in seismic zones. Ministry of Public Works and Settlement Government of Republic of Turkey. Teruna, D.R., Majid, T.A., Budiono, B., 2015. Experimental study of hysteretic steel damper for energy dissipation capacity. Adv. Civil Eng. 2015. The Japan Building Disaster Prevention Association. 2005. Standard, Guidelines, and Technical Manual for Seismic Evaluation and Seismic Retrofit of Existing Reinforced Concrete Buildings, 2001, (English version). Tokuda, Y., Taga, K., 2008. A case of structural design in which viscous dampers are used to enhance earthquake resisting performance of a building. The 14 th World Conference on Earthquake Engineering October 12 17, Beijing, China. Toranzo, L., Restrepo, J.I., Mander, J.B., Carr, A.J., 2004. Seismic design of rocking confined masonry walls with hysteretic energy dissipators and shake table validation. In: Proceedings of the 13th World Conference on Earthquake Engineering, Vancouver, Canada. Toyooka, A., 2002. Development of the inertia force driven hybrid loading system and pseudonegative stiffness control method of a MR damper. Doctor of Engineering Thesis, Kyoto University, Kyoto, Japan. Tremblay, R., and Stiemer, S.F., 1993. Energy dissipation through friction bolted connections in concentrically braced steel frames. In: ATC 17-1 Seminar on Seismic Isolation, Passive Energy Dissipation, and Active Control, 2, pp. 557 568. Tremblay, R., Degrange, G., Blouin, J., 1999. Seismic rehabilitation of a four-storey building with a stiffened bracing system. Proc. 68th Canad. Conf. Earthquake Eng. 549 554. Trombetti, T., Silvestri, S., 2004. Added viscous dampers in shear type structures: the effectiveness of mass proportional damping. J. Earthquake Eng. 8, 275 313. Trombetti, T., Silvestri, S., 2006. On the modal damping ratios of shear-type structures equipped with Rayleigh damping systems. J. Sound Vibrat. 292 (2), 21 58. Trusty, W., 2008. President, The Athena Institute. Speaker in presentation hosted by Green Building Institute: Life-cycle Assessment: From Cradle to Grave. Tsai, H.C., Lin, G.C., 1993. Optimum tuned-mass dampers for minimizing steady-state response of support-excited and damped systems. Earthquake Eng. Struct. Dyn. 22 (11), 957 973. Tsai, M.H., Chang, K.C., 2002. Higher-mode effects on the seismic response of buildings with viscoelastic dampers. Earthquake Eng. Eng. Vib. 1 (1), 119 129. Tsai, K.C., Chen, H.W., Hong, C.P., Su, Y.F., 1993. Design of steel triangular plate energy absorbers for seismic resistant construction. Earthquake Spectra 19 (3), 505 528. Tse, T.K.T., Kwok, K.C.S., Tamura, Y., 2012. Performance and cost evaluation of a smart tuned mass damper for suppressing wind-induced lateral-torsional motion of tall structures. J. Struct. Eng. 138 (4), 514 525. Tsopelas, P., Constantinou, M.C., 1994. NCEER-Taisei Corporation Research Program on Sliding Seismic Isolation Systems for Bridges: Experimental and Analytical Study of a System Consisting of Sliding Bearings and Fluid Restoring Force/Damping Devices,

1059

1060

References

Technical Report NCEER-94-0014. National Center for Earthquake Engineering Research, Buffalo, NY. Tsopelas, P., Okamoto, S., Constantinou, M.C., Ozaki, D., Fujii, S., 1994a. Isolation Systems for Bridges: Experimental and Analytical Study of Systems Consisting of Sliding Bearings, Rubber Restoring Force Devices and Fluid Dampers. Technical Report NCEER-94-002, National Center for Earthquake Engineering Research, State Univ. of New York at Buffalo, NY. Tsopelas, P.C., Constantinou, M.C., Reinhorn, A.M., 1994b. 3D-BASIS-ME computer program for nonlinear dynamic analysis of seismically isolated single and multiple structures and liquid storage tanks. Rep. No. NCEER-94-0014, National Center for Earthquake Engineering Research, State Univ. of New York at Buffalo, NY. Tsuneki, Y., Torii, S., Murakami, K., Sueoka, T., 2004b. Middle-story isolated structural system of high-rise building. The 14 th World Conference on Earthquake Engineering October 12-17, 2008, Beijing, China. Tuan, Y., Shang, G.Q., 2014. Vibration control in a 101-storey building using a tuned mass damper. J. Appl. Sci. Eng. 17 (2), 141 156. Tyler, R.G., 1977. Dynamic tests on PTFE sliding layers under earthquake conditions. Bull. N.Z. Nat. Soc. Earth. Eng. 10 (3). Tyler, R.G., 1978. Tapered steel energy dissipaters for earthquake resistant structures. Bull. N.Z. Nat. Soc. Earthquake Eng. 18 (3). Tyler, R.G., 1985. Further notes on a steel energy-absorbing element for braced framework. Bull. N.Z. Natl. Soc. Earthquake Eng. 18 (13), 270 279. September 1985. Tyler, R.G., Skinner, R.I., 1977. Testing of dampers for the base isolation of a proposed 4-storey building against earthquake attack. Proceeding 6 th Australasian on Mechanics of Structures and Materials. University of Canterbury, NZ, 376 382. UBC, 1997. Uniform Building Code. ICC. UBC, 1979. Uniform Building Code, International Conference of Building Official, Whittier, California, USA. UBC, 1982. Uniform Building Code, International Conference of Building Official, Whittier, California. UBC, 1997. Uniform Building Code, International Conference of Building Official, Whittier, California. UCD, 2001. GB 50011-2001 Code for Seismic Design of Buildings. Ministry of Construction of the People’s Republic of China and the State Quality Supervision and Quarantine Bureau. Ueng, J.-M., Lin, C.-C., Wang, J.-F., 2008. Practical design issues of tuned mass dampers for torsionally coupled buildings under earthquake loadings. Struct. Design Tall Spec. Build. 17 (1), 133 165. Uetani, K., Tsuji, M., Takewaki, I., 2003. Application of an optimum design method to practical building frames with viscous dampers and hysteretic dampers. Eng. Struct. 25 (5), 579 592. UNEP, 2007. Buildings and climate change: status, challenges and opportunities. United Nations Environment Programme, Sustainable Building and Construction Initiative 87. Uriz, P., Whittaker, A.S., 2001. Retrofit of pre-Northridge steel moment-resisting frames using fluid viscous dampers. Struct. Design Tall Build. 10 (Issue 5). Usami, T., Wang, C., Funayama, J., 2011. Low-cycle fatigue tests of a type of buckling restrained braces. Proc. Eng. 14, 956 964.

References

USRC, 2015. USRC Building Rating Systems for Earthquake Hazards: Implementation Manual. U.S. Resiliency Council. Vail, C., Hubbbell., O’Connor, B., King, J., Pall, A., 2004. Seismic Upgrade of Boeing Commercial Airplane Factory at Everett, WA, Proceedings, Thirteenth World Conference on Earthquake Engineering, Vancouver, Paper N. 3207. Van Den Berg, R.L.J., 2012. Investigation of Damping in High-Rise Buildings. Graduation Project, Tu Delft, Netherland. Vargas, R., Bruneau, M., 2007. Effect of supplemental viscous damping on the seismic response of structural systems with metallic dampers. J. Struct. Eng. 133 (10), 1434 1444. Vecchio, F.J., Collins, M.P., 1986. The modified compression-field theory for reinforced concrete elements subjected to shear. ACI Struct. J. 83 (2), 219 231. Verganelakis, V., Pall, R.T., 2004. Hightech seismic design of le nouvel Europa, Montreal. In: 13th World Conference on Earthquake Engineering, Vancouver, BC, Canada, August 1 6, Paper No. 2014. Wada, A., Huang, Y.-H., Iwata, M., 2000. Passive damping technology for buildings in Japan. Prog. Struct. Eng. Mater. 2 (3), 335 350. Wada, A., Saeki, E., Takeuchi, T., Watanabe, A., 1998. Development of unbonded brace. Nippon Steel’s Unbonded Braces (promotional document) 1 16. Wakabayashi, M., Nakamura, T., Katagihara, A., Yogoyama, H., Morisono, T., 1973. Experimental Study on the Elasto-Plastic Behaviour of Braces enclosed by Precast Concrete Panels under Horizontal Cyclic Loading 10, 1041 1044. Wakabayashi, M., Nakamura, T., Iwai, S., Hayashi, Y., 1984. Effects of Strain Rate on the Behaviour of Structural Members. Proceedings 8th World Conference on Earthquake Engineering, San Francisco, Ca, IV, 491 498. Wang, S.G., Du, D.S., Liu, W.Q., 2009. Research on key issues about seismic isolation design of high-rise buildings structure. In: Proceedings of the 11thWorld Conference on Seismic Isolation, Energy Dissipation and Active Vibration Control of Structures, pp. 17 21, Guangzhou, China. Wang, D., Tse, T.K.T., Zhou, Y., Li, Q., 2015. Structural performance and cost analysis of wind-induced vibration control schemes for a real super-tall building. Struct. Infrastruct. Eng. Maint. Manage. Life-Cycle Design Perfor. 11 (8), 990 1011. Wang, L., Zhao, X., Zheng, Y.M., 2016. A combined tuned damper and an optimal design method for wind-induced vibration control for super tall buildings. Struct. Design Tall Spec. Build. 25, 468 502. Wanitkorkul, A., Filiatrault, A., 2008. Influence of passive supplemental damping systems on structural and nonstructural seismic fragilities of a steel building. Eng. Struct. 30 (3), 675 682. Warburton, G.B., 1981. Optimum absorber parameters for minimizing vibration response. Earthquake Eng. Struct. Dyn. 9 (3), 251 262. Warburton, G.B., 1982. Optimum absorber parameters for various combinations of response and excitation parameters. Earthquake Eng. Struct. Dyn. 10 (3), 381 401. Warburton, G.B., Ayorinde, E.O., 1980. Optimum absorber parameters for simple systems. Earthquake Eng. Struct. Dyn. 8, 197 217. Warnitchai, P., Pinkaew, T., 1998. Modelling of liquid sloshing in a rectangular tank with flow-dampening devices. Eng. Struct. 20 (7), 593 600.

1061

1062

References

Watanabe, A., Hitomi, Y., Saeki, E., Wada, A., Fujimoto, M., 1988. Properties of brace encased in buckling-restraining concrete and steel tube. Proc. Ninth World Conf. Earthquake Eng. IV, 719 724. Webster, M., 2001. Role of the structural engineer in sustainable design. Proceeding of the First International Conference on Ecological Building Structure, CRC Press, China. Wen, P., Baifeng, S., 2008. Two step design method for base isolation structures. In: The 14 th World Conference on Earthquake Engineering, October 12 17, Beijing, China. Wen, Y.K., 1976. Method for random vibration of hysteretic systems. J. Eng. Mech. 102 (2), 249 263. Weng, D.G., Zhang, C., Lu, X.L., Zeng, S., Zhang, S.M., 2012. A simplified design procedure for seismic retrofit of earthquake-damaged RC frames with viscous dampers. Struct. Eng. Mech. 44 (5), 611 631. Wenzel, H., Pichler, D., Schedler, R., 1991. Ambiente Schwingungsmessungen zur System und Schadenserkennung an Tragwerken. Bauingenieur 74 (3), 115 123 (in German). Wereley, N.M., Pang, L., Kamath, G.M., 1998. Idealized hysteresis modeling of electrorheological and magnetorheological dampers. J. Intelligent Mater. Syst. Struct. 9 (8), 642 649. Whittaker, A.S., Bertero, V.V., Thompson, C.I., Alsonson, L.J., 1991. Seismic testing of steel plate energy dissipation devices. Earthquake Spectra 7 (4), 563 604. Whittle, J.K., Williams, M.S., Karavasilis, T.L., Blakeborough, A., 2012. A comparison of viscous damper placement methods for improving seismic building design. J. Earthquake Eng. 16 (4), 540 560. Wienser, K.B., 1979. Tuned Mass Dampers to Reduce Building Wind Motion. Preprint, 3510. ASCE Convention and Exposition, Boston, MA, pp. 1 21. Wigle, V.R., Fahnestock, L.A., 2010. Buckling-restrained braced frame connection performance. J. Construct. Steel Res. 66, 65 74. Wijanto, S., 2012. Generic buckling restrained brace systems. Master Thesis. The University of Auckland, Auckland, New Zealand. Willard, B., 2002. The Sustainability Advantage Seven Business Case Benefits of a Triple Bottom Line. New Society, British Columbia. Williams, M.S., Blakeborough, A., 2001. Laboratory testing of structures under dynamic loads: an introductory review. R. Soc. Lond. 359, 1651 1669. Wirsching, P.H., Yao, J.T.P., 1973. Safety design concepts for seismic structures. Comput Struct. 3, 809 826. Wolff, E.D., Ipek, C., Constantinou, M.C., Tapan, M., 2014. Effect of viscous damping devices on the response of seismically isolated structures. Earthquake Eng. Struct. Dyn. 44 (2), 185 198. Wolfram Research, Inc., 2016. Mathematica, Version 10.4, Champaign, IL. Won, A.Y.J., Pires, J.A., Haroun, M.A., 1996. Stochastic seismic performance evaluation of tuned liquid column dampers. Earthquake Eng. Struct. Dyn. 25, 1259 1274. Wongprasert, N., Symans, M.D., 2005. Experimental evaluation of adaptive elastomeric baseisolated structures using variable-orifice fluid dampers. J. Struct. Eng. 131 (6), 867 877. Wu, J.C., Chang, C.H., 2006. Design table of optimal parameters for tuned liquid column damper responding to earthquake. In: Proceedings 4th international conference on earthquake engineering, Taipei, 12 13 October.

References

Wu, J.-C., Shih, M.-H., Lin, Y.-Y., Shen, Y.-C., 2005. Design guidelines for tuned liquid column damper for structures responding to wind. Eng. Struct. 27, 1893 1905. Wu, J.C., Chang, C.H., Lin, Y.Y., 2009. Optimal design of non-uniform tuned liquid column dampers in horizontal motion. Sound Vib. 326 (1), 104 122. Wyatt, T.A., 1977. Mechanisms of damping. proceedings of a symposium on dynamic behaviour of bridges. Transport and Road Research Laboratory, Crowthorne. Xia, C., Hanson, R.D., 1992. Influence of ADAS element parameters on building seismic response. J. Struct. Eng. 118 (7), 1903 1918. Xia, Y., Ni, Y.Q., Zheng, P., Liao, W.Y., Ko, J.M., 2011. Stress development of a supertall structure during construction: field monitoring and numerical analysis. Comput. Aided Civil Infrastruct. Eng. 26, 542 559. Xia, Y., Zheng, P., Ni, Y.Q., Zhu, H.P., 2014. Deformation monitoring of a super-tall structure using real-time strain data. J. Eng. Struct. 67, 29 38. Xie, Q., 2005. State of the art of buckling-restrained braces in Asia. J. Construct. Steel Res. 61, 727 748. Xu, K., Igusa, T., 1992. Dynamic characteristics of multiple substructures with closely spaced frequencies. Earthquake Eng. Struct. Dyn. 21 (12), 1059 1070. Xu, Y., Ng, C., 2008. Seismic protection of a building complex using variable friction damper: experimental investigation. J. Eng. Mech. 134 (8), 637 649. Xu, Y.L., 1996. Parametric study of active mass dampers for wind-excited tall buildings. Eng. Struct. 18 (1), 64 76. Xu, Y.L., Chen, B., 2008. Integrated vibration control and health monitoring of building structures using semi-active friction dampers: Part I—methodology. Eng. Struct. 30 (7), 1789 1801. Xu, Y.L., Samali, B., Kwok, K.C.S., 1992. Control of along-wind response of structures by mass and liquid dampers. J. Eng. Mech. 118 (1), 20 39. Xu, Z.-D., Shi, C.-F., 2008. Energy dissipation analysis on real application by using viscoelastic dampers. In: The 14 th World Conference on Earthquake Engineering, October 12 17, Beijing, China. Xu, Z.-D., Sha, L.-F., Zhang, X.-C., Ye, H.-H., 2012. Design, performance test and analysis on magnetorheological damper for earthquake mitigation. Struct. Control Health Monit. 20 (6), 956 970. Xue, M., Lu, L.W., 1994. Interaction of infilled steel shear wall panels with surroundings frame members. In: Proceedings of Structural Stability Council Annual Technical Session, SSRC, Bethlehem, PA, pp. 339 354. Yalla, S., Kareem, A., 2003. Semiactive tuned liquid column dampers: experimental study. J. Struct. Eng. 129 (7), 960. 960 971. Yalla, S.K., 2001. Liquid Dampers for Mitigation of Structural Response: Theoretical Development and Experimental Validation. Ph.D. Thesis, Notre Dame, Indiana. Yalla, S.K., Kareem, A., 2000. Optimum absorber parameters for tuned liquid column dampers. J. Struct. Eng. 126 (8), 906 915. Yalla, S.K., Kareem, A., Kantor, J.C., 2001. Semi-active tuned liquid column dampers for vibration control of structures. Eng. Struct. 23, 1469 1479. Yamaguchi, H., Harnpornchai, N., 1993. Fundamental characteristics of multiple tuned mass dampers for suppressing harmonically forced oscillations. Earthquake Eng. Struct. Dyn. 22 (1), 51 62.

1063

1064

References

Yamamoto, M., Sone, T., 2014a. Damping systems that are effective over a wide range of displacement amplitudes using metallic yielding component and viscoelastic damper in series. Earthquake Eng. Struct. Dyn. 43 (14), 2097 2114. Yamamoto, M., Sone, T., 2014b. Behavior of active mass damper (AMD) installed in highrise building during 2011 earthquake off Pacific coast of Tohoku and verification of regenerating system of AMD based on monitoring. Struct. Control Health Monit. 21 (4), 634 647. Yamamoto, M., Aizawa, S., Higashino, M., Toyama, K., 2001. Practical applications of active mass dampers with hydraulic actuator. Earthquake Eng. Struct. Dyn. 30 (11), 1697 1717. Yan, N., Wang, C.M., Balendra, T., 1998. Composite mass dampers for vibration control of wind-excited towers. J. Sound Vib. 213 (2), 301 316. Yan, N., Wang, C.M., Balendra, T., 1999. Optimal damper characteristics of ATMD for buildings under wind loads. J. Struct. Eng. 125 (12), 1376 1383. Yang, G., Ramallo, J.C., Spencer, B.F., Jr., Carlson, J.D., Sain, M.K., 2000. Large-scale MR fluid dampers: dynamic performance considerations. In: Proceedings of International Conference on Advances in Structure Dynamics, Hong Kong, pp. 341 348. Yang, G., Jung, H.J., Spencer, B.F., Jr. 2001. Dynamic modeling of full-scale MR dampers for civil engineering applications. In: Proc., U.S. Japan Workshop on Smart Structures for Improved Seismic Performance in Urban Region, Seattle, pp. 213 224. Yang, J.N., Lin, S., Kim, J.H., Agrawal, A.K., 2002a. Optimal design of passive energy dissipation systems based on H1 infinity and H2 performances. Earthquake Eng. Struct. Dyn. 31 (4), 921 936. Yang, G., Spencer Jr., B.F., Carlson, J.D., Sain, M.K., 2002b. Large-scale MR fluid dampers: modeling, and dynamic performance considerations. Eng. Struct. 24, 309 323. Yang, G., Spencer, B.F., Jung, H.-J., Carlson, J.D., 2004. Dynamic modeling of large-scale magnetorheological damper systems for civil engineering applications. J. Eng. Mech. 130, 1107 1114. Yang, D., Lu, Z., Zhu, H., Li, Z., 2014. Simplified design method for shear-valve magnetorheological dampers. Earthquake Eng. Eng. Vib. 13, 637 652. Yang, Y., Li, Q.S., Yan, B.W., 2017. Specifications and applications of the technical code for monitoring of building and bridge structures in China. Adv. Mech. Eng. 9 (1), 1 10. Yao, J.T.P., 1972. Concept of structural control. J. Struct. Div. ASCE 98 (7), 1567 1574. Yokoo, Y., Akiyama, H., 1972. Lateral vibration and damping due to wind and earthquake effects. In: Proceedings of International Conference on Planning and Design of Tall Buildings, Vol. 11 17, ASCE, NY. Yokota, H., Saruta, M., Nakamura, Y., Satake, N., Okada, K., 1992. Structural control for seismic loads using viscoelastic dampers. In: Earthquake Engineering 10 th Conference, Balkema, Rotterdam, pp. 2441 2446. Yoon, S.W., Ju, Y.K., 2004. Dynamic properties of tall buildings in Korea. Proceedings of CTBUH 2004, (CD-ROM), Council of Tall Buildings and Urban Habitat, Chicago. Yoshida, O., Dyke, S.J., 2005. Response control of full-scale irregular buildings using magnetorheological dampers. J. Struct. Eng. 131 (5), 734 742. Yu, J., Wakahara, T., Reed, D.A., 1999. A nonlinear numerical model of the tuned liquid damper. Earthquake Eng. Struct. Dyn. 28, 671 686.

References

Yu, J.-K., 1997. Nonlinear Characteristics of Tuned Liquid Dampers. PhD Thesis, University of Washington, United States. Zahrai, S.M., Ghannadi-Asl, A., 2008. Seismic performance of TMDs in improving the response of MRF buildings. Sci. Iranica 15 (1), 21 33. Zhang, Z., Cho, C., 2009. Experimental study on damping ratios of in-situ buildings. World Acad. Sci. Eng. Technol. 26, 614 618. Zhang, C., Zhang, Z., Zhang, Q., 2012. Static and dynamic cyclic performance of a lowyield-strength steel shear panel damper. J. Construct. Steel Res. 79, 195 203. Zhao, X., Han, H., 2016. Integrated optimal seismic design of mega-brace da Leonard mpers under story drift constraint for super tall buildings. In: The 2016 Structures Congress, 28 th August-1 st September, Jeju Island, Korea. Zhou, F., Tan, P., Yan, W., Wei, L., 2002. Theoretical and experimental research on new system of semi active control. Earthquake Eng. Struct. Dyn. 1 (1), 130 136. Zhou, F.L., 2001. Seismic isolation of civil buildings in the People’s Republic of China. Prog. Struct. Eng. Mater. 3, 268 276. Zhou, Y., Lu, X., Weng, D., Zhang, R., 2012. A practical design method for reinforced concrete structures with viscous dampers. Eng. Struct. 39, 187 198. Zimmer, M., 1999. Characterization of Viscoelastic Materials for use in Seismic Energy Dissipation Systems. Master of Science Thesis, Department of Civil, Structural and Environmental Engineering, University at Buffalo, State University of New York, Buffalo, NY. Zuk, W., 1968. Kinetic structures. Civil Eng. 39 (12), 62 64.

1065

Index Note: Page numbers followed by “f” and “t” refer to figures and tables, respectively.

A A&H Costum, 206 Aanti-yaw@ mechanism, 794 795 Abeno Harukas 300 project, 176, 177f, 962t Absolute-displacement damper designs, 78 Acceleration criteria, 319 320, 346 feedback, 381 Accelerometers, 740 742, 816 Acceptance criteria, 12, 414 417 component force-deformation curve, 414f components force-and deformation-controlled actions, 415t force-deformation modeling and acceptance criteria, 416f load combinations, 416 417 nonlinear procedure component capacity, 417t inspection, 476 477 ACT. See Active control theory (ACT) Action, 12 Active control parameter, 380 strategy, 453 systems, 217, 218f, 219 220, 517 Active control theory (ACT), 366, 376 Active damping systems, 492 493, 497t Active fault, 12 Active mass dampers (AMD), 9, 486, 510 Active mass driver system (AMD system), 365, 370 371 Active systems, 217 232, 459, 499 500, 517 519, 926 active mass damper, 510 active tuned mass damper, 510 512, 511f actuators, 223 224 ATMD, 222 223, 229 230 case studies, 534, 904 919 Thyssenkrup Test Tower, Rottweil, Germany, 904 919 control strategies, 230 231 dampers manufactures, 232 future directions, 231 232 HMD, 512 513 interaction of active damping system and DSF systems, 460f literature review, 365 367 semiactive, hybrid dampers manufactures, 232

step-by-step procedure, 367 400, 368f bandwidth specification of active and semiactive systems, 370 build simplified and reduced model, 372 376 building and site categorization, 369 building fundamental properties and preliminary structural analyses, 369 370 damper type, configuration, and distribution of sensors and actuators, 370 372 design of controller system, 372 376 selecting structural system, 369 selection of performance objective, 370 verification analyses and quality control, 394 400 Active tendon mechanisms (ATM), 372 Active tendons systems, 224 control of cable-stayed bridges, 224f Active tuned mass damper (ATMD), 220, 221f, 222 223, 229 230, 365, 485, 510 512, 511f adaptive radius TMD systems, 230f dampers, 395 397 design, 377 383 Active variable stiffness devices (AVSD), 9 Active viscous damping system (AVDS), 366 Active-type damper, 938 941 Actuators, 223 224, 223f, 459 Adaptive systems, 221 Adaptive/smart structures, 217 Added components, 239 240 Added damping-added stiffness systems (ADAS systems), 152 155, 152f metallic dampers optimum geometry, 153f Additional mass systems, 8 9 Additional/Added/Supplementary damping, 13 Adequate corrosion protection, 134 Aerodynamic damping, 55 Aeroelastic damping, 13 model, 744 wind tunnel model, 795 796 AGOM, 133 AIJ. See Architectural Institute of Japan (AIJ) Air-handling units, 32 Algorithm-based optimization procedures, 430 432 control theory-based algorithms, 432 GAs, 430 432

1067

1068 Index

Algorithm-based optimization procedures (Continued) gradient-based algorithms, 430 heuristic algorithms, 432 Allianz Tower, Milan, Italy building overview, 717f comparison different design strategies cost benefit analysis, 732 damping solution consideration, 732 structural options consideration, 732 damping overview architectural integration strategy, 730 damping strategy, 725, 725f, 726f damping type, 726 727 elevator and other mechanical devices integration strategy, 730 experimental tests, 730 731 MEP integration strategy, 730 monitoring system, 731 structural and damping design, 727 730, 728f, 729f introduction/history, 718 lesson learned and recommendations design innovative solutions, 733 difficulties in design, 732 733 improvements, 733 project data, 716 718 structural system, 718 725, 718f additional damping provided by damping system, 720 722 building code, 723 724 building cost vs. damping cost, 722 723 building fundamental periods, 719 720 damping strategy utilization, 720 peer-reviewed project, 725 Allowable story drift, 319, 319t Allowable stress design, load combinations for, 283 Aluminum, 149 151 alloys, 150 151 AMD. See Active mass dampers (AMD) AMD system. See Active mass driver system (AMD system) American Society of Civil Engineers (ASCE), 81 Amplitude amplitude-dependent regions, 62 63 scaling, 274 Analysis procedure selection, 288 298, 289t, 327 nonlinear procedure, 297 298 response spectrum, 289 297 Analytical and numerical fluid models, 202 203 ANPSD Method. See Averaged Normalized Power Spectral Density Method (ANPSD Method) Anti-seismic devices, 478 479

Architectural aspects, 438 450 base isolation systems, 447 450 distributed damping systems, 439 443 mass damping systems, 444 446 components, 439 Architectural Institute of Japan (AIJ), 82 84, 508 Architecturally integrated building systems, 438 AS/NZS. See Australia and New Zealand wind standard (AS/NZS) ASCE. See American Society of Civil Engineers (ASCE) ASCE 7 16, 263 264, 292 294, 347 348 code, 69 requirements, 468 ASCE 41 13 code, 69 Aspect ratio, 12 ATM. See Active tendon mechanisms (ATM) ATMD. See Active tuned mass damper (ATMD) Atushi Building, Xin Jiang, China building overview, 669f comparison different design strategies cost benefit analysis, 677 damping solution consideration, 677 structural options consideration, 677 damping overview architectural integration strategy, 676 damping strategy, 672 673 damping type, 673 elevator and other mechanical devices integration strategy, 676 experimental tests, 676 MEP integration strategy, 676 monitoring system, 676 677 structural and damping design, 673 675 damping ratio result summary, 674t fluid viscous damper drawing, 673f introduction/history, 669 lesson learned and recommendations design innovative solutions, 677 difficulties in design, 677 improvements, 677 project data, 668 structural system, 669 672, 671f additional damping provided by damping system, 671 building code, 672 building cost vs. damping cost, 671 building fundamental periods, 670, 670f damping strategy utilization, 670 peer-reviewed project, 671f, 672, 672f structure dynamic periods, 670t

Index 1069

toggle brace damper in X direction, 671f in Y direction, 672f Austenite, 151 Australia and New Zealand wind standard (AS/NZS), 66, 84 ¨ NORM B4014 Teil 1, 72 Austria O Automotive sector, 925 Auxiliary damping device, 438 AVDS. See Active viscous damping system (AVDS) Averaged Normalized Power Spectral Density Method (ANPSD Method), 914 AVSD. See Active variable stiffness devices (AVSD) Axial shortening action, 58 60

B B. Don, 134 Backstay effects, 12, 261 262 Bandwidth frequency, 1001 1002 Barber and Wen model, 155 Base isolated buildings, 244 245 Base isolation systems, 207 217, 207f, 447 450, 462, 472, 486 491, 500, 505, 513, 518 519, 534, 888 904, 1009 1012. See also Distributed damping systems; Mass damping systems application, 208 210 devices, 217 manufactures, 215 217 standard code requirements, 486 487 step-by-step procedure for, 347 365 building and site categorization, 349 default upper and lower bound multipliers for isolation bearings, 362t design procedures literature review, 347 349 estimated limits of aspect ratio, 350f isolation system modeling, 360 363 isolation system primary design, 357 359 isolation type and distribution, 351 353 nonlinear time history analysis, 360 plan distribution example for Nunoa Capital Building, 354f preliminary structural analysis, 354 357 quality assurance, 364 365 response spectrum analysis, 360 structural system(s) selection, 349 351 target vibration period and total damping ratio selection, 351 updating building model and final analyses, 359 364

usual practice, 363 364 structure testing, 490 491 testing, 487 490 complete unit tests, 490 process, 497t prototype tests, 488 489 qualification tests, 489 quality control procedures, 490 system characterization tests, 487 types, 210 215 advantages and disadvantages of isolation devices, 211t bearing systems, 212 213 seismic isolation types, 212t sliding systems, 213 215 Nunoa Capital Building, Santiago, Chile, 888 904 view of isolation floor based on ground and rubber bearings, 448f Base level, 13 Base shear, 13 Basic PO Equivalent to New Building Standards (BPON), 406 Basic POs for existing building (BPOE), 405 Basic Safety Earthquake-1 (BSE-1E), 407 Basic Safety Earthquake-2 (BSE-2E), 407 Bauschinger curve, 146 Bauschinger effect, 146 Beam, 13 beam-type damper, 439, 440f Bearing systems, 210, 212 213 rubber bearing and force-deformation behavior, 212f Beijing Yintai Center, Beijing, China acceleration of stories under wind load, 588t building overview, 580f comparison different design strategies cost benefit analysis, 590 damping solution consideration, 590 structural options consideration, 590 damper layout on elevation, 584f layout on floor plan, 583f damping overview architectural integration strategy, 589 damping strategy, 585 damping type, 585 elevator and other mechanical devices integration strategy, 589 experimental tests, 588t, 589 maintenance strategy, 589 590 MEP integration strategy, 589 monitoring system, 589

1070 Index

Beijing Yintai Center, Beijing, China (Continued) structural and damping design, 585 588 frame support and standard layout structural plan, 582f fundamental frequencies and mode shapes, 582f introduction/history, 581 lesson learned and recommendations design innovative solutions, 591 difficulties in design, 591 improvements, 592 project data, 577 581 structural effect under frequent earthquake, 588t under rare earthquake, 589t structural system, 581 585 additional damping provided by damping system, 582 583 building code, 584 585 building cost vs. damping cost, 583 building fundamental periods, 581 damping strategy utilization, 581 peer-reviewed project, 585 viscous damper composition diagram, 586f viscous damper detail photo, 586f Betchel Power Corporation, 152 Bilinear hysteretic simplification, 979 model, 143, 143f BIM. See Building information modeling (BIM) Bitumen compound, 124 Bloomberg Tower, New York city, USA, 200, 809f comparison different design strategies cost benefit analysis, 818 damping solution consideration, 818 structural options consideration, 817 damping overview architectural integration strategy, 815 elevator and mechanical devices integration strategy, 815 experimental tests, 815 maintenance strategy, 817 MEP integration strategy, 815 monitoring system, 816 strategy, 811 812 structural and damping design, 813 815 type, 812 introduction/history, 810 lesson learned and recommendations design innovative solutions, 818 difficulties in design, 818 improvements, 818 pictures of installed TMD, 813f project data, 808 810

structural system, 810 811 additional damping provided by damping system, 811 building code, 811 building cost vs. damping cost, 811 building fundamental periods, 811 damping strategy utilization, 811 design forces, 811 peer-reviewed project, 811 TMD general scheme, 813f typical FFT magnitude and phase of x1 acceleration measurement, 817f typical sample of time history acceleration measurements, 816f wind tunnel test, 812f BLWTL. See Boundary Layer Wind Tunnel Laboratory (BLWTL) BMT fluid mechanics, 78 85 Bouc Wen model, 143, 144f Boundary conditions and common assumptions, 261 264 effective stiffness for RC components, 262 263 modification factors of damper properties, 263 264 SSI, 261 262 Boundary Layer Wind Tunnel Laboratory (BLWTL), 776 BPOE. See Basic POs for existing building (BPOE) BPON. See Basic PO Equivalent to New Building Standards (BPON) Brace stiffness, 311 312 Braced moment frames, 540 Bracket-type systems, 439, 440f BRAD. See Buckling-restrained axial damper (BRAD) Brant Hydraulics Corporation, 133 BRBs. See Buckling restrained bracings (BRBs) Bridgestone, 215 British Standard BS ISO 4866:2010, 72 BSE-1E. See Basic Safety Earthquake-1 (BSE-1E) BSL. See Japanese Building Standard Law (BSL) Buckling restrained bracings (BRBs), 186 191, 401, 504, 551 552, 572 behavior, 188 189, 189f connection type, 192f, 193t cross section, 190f damper, 471 displacement between types of BRB bracing configurations, 189f geometry and typical bracing layout, 188f knee braces scheme, 190f Buckling-restrained axial damper (BRAD), 194 195

Index 1071

Building health monitoring, 505 513, 923 active, semiactive and hybrid damping systems, 507 508 base isolation systems, 508 case study examples, 508 513 active, semiactive and hybrid damped systems, 510 513 base isolation systems, 513 distributed damped systems, 509 isolated damped systems, 510 distributed damping systems, 507 isolated damping systems, 507 Building information modeling (BIM), 924 Building(s), 13, 949 950 code(s), 300 301, 737 requirements, 466 467 cost, 737 damping, 7 8 deformation, 90 drift ratio, 13 with energy dissipation devices, 245 evaluation process, 402 403 function, 958 961, 960f fundamental periods, 736 general trends for, 943 945 buildings over 200 m in United States, 951f integration modes, 438 motion, 109 performance targets, 437 438 response principles, 41 54 MDOF system, 47 53 SDOF system, 41 47 stiffness, 7 8 structure dampers testing, 480 481 testing tools for damper attached with, 481 summarized data of buildings with damper, 961, 962t system interaction architectural aspects, 438 450 elevators, 450 454 fac¸ade, 457 464 mechanical systems, 455 457 TMD, 770 771 wind vibration, 79 80 Buildings over 200 m in United States (Database 2), 943 948. See also Damping system(s) general trends for buildings, 943 945 general trends for damping system, 945 947 trends for damping system category, 947 trends for damping system type, 948 Bulged shaft, 167 Buttress with viscous dampers, 443, 444f

C California Building Code (CBC), 597 California Strong Motion Instrumentation Program (CSMIP), 608, 922 923 Canadian building code, 73 Cantilever action, 58 60, 60f Capacity design, 13 reduction factor, 4 Care of works of damping systems, 526 Cast Connex, 194 CBC. See California Building Code (CBC) CBD. See Central Business District (CBD) CBF. See Concentrically braced frame (CBF) CEEDI. See China Electronics Engineering Design Institute (CEEDI) Central Business District (CBD), 581 Certified Rating Professional (CRP), 530 CFD. See Computational fluid dynamics (CFD) Characteristic Earthquake, 13 Characteristic strength, 311, 358 Chevron dampers, 441 442, 443f Chevron frames, 116 dampers in chevron-braced frames, 116f Chiba Port Tower, 198 Chile earthquake, 890 Chilean isolation code, 898 899 Chilean seismic design standards, 890 Chilean standard NCh2745, 900 901 China Electronics Engineering Design Institute (CEEDI), 577 Chinese code, 243 245, 471 base isolated buildings, 244 245 buildings with energy dissipation devices, 245 GB 50011, 347 348 Chinese standard, 502, 504 code, 68 testing, 478 Circular plate damper (CPD), 160, 161f Circular tank, 339 Circular/torsional TLCDs, 205 Citicorp Building, New York, USA, 734f, 736f, 737f, 738f comparison different design strategies, 744 745 cost benefit analysis, 745 damping solution consideration, 745 structural options consideration, 744 damping overview architectural integration strategy, 743 development process, 744 elevator and mechanical devices integration strategy, 743 experimental tests, 744

1072 Index

Citicorp Building, New York, USA (Continued) maintenance strategy, 744 MEP integration strategy, 743 monitoring system, 744 strategy, 740, 740f structural and damping design, 742 743 type, 740 742, 741f introduction/history, 735 lesson learned and recommendations design innovative solutions, 745 difficulties in design, 745 improvements, 745 project data, 734 735 structural system, 735 740 additional damping provided by damping system, 737 building code, 737 building cost vs. damping cost, 737 building fundamental periods, 736 damping strategy utilization, 736 design forces, 738 739 expected performance, 740 lower floors structure under construction, 739f peer-reviewed project, 738 upper floors structure under construction, 739f Citicorp Center, 198 Clamp friction force, 171 Clamped RB models, 260 261 Classical linear modal analysis with eigenvectors, 259 260 Classical Maxwell model, 425 Closed-loop complete-feedback control algorithm, 365 Closed-loop control system. See Feedback control Code perception criteria, 81 84 Code prescriptive procedures, 267 Code seismic deflection limits, 90, 91t Code wind deflection limits, 86t, 95 Cold War technology, 135 136 “Collapse prevention”, 18, 925 Collocated control system, 371 372 Columbia Seafirst Center, 539f, 540, 541f, 542f Columbia Tower, Seattle, USA case study with dynamic modification summary, 536t comparison different design strategies, 546 cost benefit analysis, 546 damping solution consideration, 546 structural options consideration, 546 damping overview architectural integration strategy, 545 damping strategy, 543 544

damping type, 545 elevator and other mechanical devices integration strategy, 545 experimental tests, 546 maintenance strategy, 546 MEP integration strategy, 545 monitoring system, 546 structural and damping design, 545 introduction/history, 540 lesson learned and recommendations design innovative solutions, 546 difficulties in design, 546 possible improvements, 547 project data, 534 540 structural system, 540 542 additional damping provided by damping system, 541 building code, 542 building cost vs. damping cost, 542 building fundamental periods, 540 541 Columbia Center braced moment frame system, 543f damping strategy utilizing, 541 peer-reviewed project, 542 Column-type damper, 439, 440f Comcast Center, Philadelphia, USA, 784f comparison different design strategies cost benefit analysis, 788 damping solution consideration, 787 structural options consideration, 787 damping overview, 786 787 architectural integration strategy, 787 elevator and mechanical devices integration strategy, 787 experimental tests, 787 maintenance strategy, 787 MEP integration strategy, 787 monitoring system, 787 strategy, 786 structural and damping design, 786 787 type, 786 introduction/history, 783 lesson learned and recommendations design innovative solutions, 788 difficulties in design, 788 improvements, 788 project data, 783 structural system, 783 786 additional damping provided by damping system, 785 building code, 786 building cost vs. damping cost, 785 786 building fundamental periods, 785 damping strategy utilization, 785

Index 1073

expected performance, 786 peer-reviewed project, 786 Commissioning and system tuning, 493 500 base isolation systems, 500 distributed damping systems, 494 495 isolated damping systems, 495 500 Complete quadratic combination method (CQC method), 253, 293 294, 997 Complete unit tests, 490 Completion inspection, 476 477 Composite, 953 Composite Concrete Core, 865 Computational fluid dynamics (CFD), 105 106 Computational structural analysis modeling, 252 Concealed dampers, 439, 441 443 Concealed distributed damping systems. See Exposed distributed damping systems Concentrically braced frame (CBF), 16, 164 Concrete, 810 811, 953 Connor Tower, Manila, Philippines building overview, 702f comparison different design strategies cost benefit analysis, 714 716, 716t damping solution consideration, 714 structural options consideration, 714 damping overview architectural integration strategy, 712 damping strategy, 709 710 damping type, 710 development process, 713 elevator and other mechanical devices integration strategy, 712 experimental tests, 713 maintenance strategy, 713 MEP integration strategy, 712 monitoring system, 713 structural and damping design, 711 712 introduction/history, 703 704 lesson learned and recommendations design innovative solutions, 716 difficulties in design, 716 improvements, 716 project data, 701 703 structural system, 704 709 additional damping provided by damping system, 705 building code, 706 707 building cost vs. damping cost, 705 706 building fundamental periods, 704, 705f damping strategy utilization, 704 705 design forces, 707 peer-reviewed project, 707 performance, 707 709 3D structural model, 706f

Conservatory structural system, 820 Constricted tube, 167 Construction inspection, 476 477 Construction phase VDD prototype tests, 606 608 Control algorithms, 218, 219f force vector, 219 methods, 218 219 process, 219 220 strategies, 230 231 system, 218 219 theory-based algorithms, 430, 432 Controller system, 372 376 design of ATMD, 377 383 design of semiactive systems, 383 394 Conventional mass damper. See Single mass damper Conventional TMDs. See Single TMDs Convergence studies, 259 Copper, 148 cyclic stress strain relationship for copper damper, 150f stress strain relationship, 149f Core systems, 958 CoreBrace, 191, 194 Costums Residential, Auckland, New Zealand building elevation with distributed damping system, 689f building overview, 678f comparison different design strategies cost benefit analysis, 699 700 damping solution consideration, 699 structural options consideration, 699 damping overview architectural integration strategy, 696 damping strategy, 686 688 damping type, 688 692 elevator and other mechanical devices integration strategy, 697 experimental tests, 697, 697t maintenance strategy, 699 MEP integration strategy, 697 monitoring system, 697 698 structural and damping design, 692 696 introduction/history, 679 680 lesson learned and recommendations design innovative solutions, 701 difficulties in design, 701 improvements, 701 project data, 678 679 proposed TMD system longitudinal view, 690f transverse view, 690f structural system

1074 Index

Costums Residential, Auckland, New Zealand (Continued) additional damping provided by damping system, 684 685, 685t building code, 685 building cost vs. damping cost, 685 building fundamental periods, 681, 682f damping strategy utilization, 681 684 design forces, 685 gravity system, 680 681 maximum acceleration criteria, 686t peer-reviewed project, 685 performance, 685 stability, 680, 681f upper floors elevation view, 691f Coulomb dampers. See Friction dampers Coulomb damping, 13 Coulomb friction, 170, 981 SDOF model, 171f Coulomb’s law, 58 60 Coupled-two-beams (CTB), 260 261 Coupling beam, 13 dampers, 124 125, 125f CPD. See Circular plate damper (CPD) CQC method. See Complete quadratic combination method (CQC method) Cracking modification factors, 262 Critical damping, 13 “Critical excitation”, 102 103 “Critical tip deflection”, 61 Cross-layer connection, 120 121 interconnection, 120 121 Cross-sectional area of tube, 343 344 Crossover working care, 526 CRP. See Certified Rating Professional (CRP) Crystal Tower, 200, 200f CSMIP. See California Strong Motion Instrumentation Program (CSMIP) CTB. See Coupled-two-beams (CTB) CTBUH, 949 950, 953, 958 961 CW2012, 348 349 Cycle testing, 606 608

D d’Alembert principle, 41 Damaged-oriented LCA, 97 101 Damped link element, 119 120, 120f Damped MDOF system, 51 Damped outrigger system, 443, 457 and building core, 445f Damped structural model construction, 312 317

effects of variation in damper properties, 315 316 force-velocity relationship for nonlinear viscous damper, 316f friction and hysteretic dampers modeling, 314 315 load combinations, 316 317 viscous/viscoelastic dampers modeling, 314, 314f Damped structure, 985 Damper(s), 4, 14, 109, 111, 119 120, 128, 240, 994 in chevron-braced frames, 116f configurations, 439 criteria, 320 321, 346 damper/isolation, 20 in diagonal-braced frames, 115f distribution, 300 301, 329 331 fluid viscous, 971 975 friction, 981 983 geometrical configuration, 299 hysteretic, 978 981 manufacturer, 251 in mega brace frames, 120f modeling description, 139 140 modification factors of damper properties, 263 264 performance curves for, 242f placement, 439 power, 306 stiffness, 156 160 and system design procedure, 243f system from concept to production process, 248 249 in tall buildings, 1, 5 comparison, 232 TLCD, 1008 1009 TLD, 1005 1007 TMD, 983 1005 in toggle-braced frames, 117f type, 930 configuration, and distribution, 298 302 selection, 328 329 viscoelastic, 975 978 Damping, 13 14, 45 46, 921, 1005 alternative devices, 103 106 alternative to damping devices, 103 106 category, 930 coefficient, 302 306, 389 390 of TMDs, 333 334 considerations, 39 constant, 13 cost, 700, 737 effect on building, 73 78

Index 1075

devices, 11, 28 29, 232, 431 432 characterization, 248 industrial application, 30 37 properties, 251 technology, 107 effect on building, 73 78 development considerations, 77 78 elements, 35 environmental and economic considerations, 90 102 factor, 290, 290t height vs., 949 953 damping trends in tallest buildings, 954f height criteria for damping categories, 958f world distribution of top 20 tallest buildings with damping systems, 958f matrix, 51 52, 55 56 modeling, 55 57 preliminary design, 302 312 constitutive relation curves of different damping index, 303f displacement-dependent dampers properties estimation, 308 312 horizontal magnification factors, 304t vertical magnification factors, 305t viscoelastic dampers property estimation, 306 308 viscous dampers properties estimation, 302 306 principles of building response, 41 54 equivalent viscous damping, 53 54 MDOF system, 47 53 SDOF system, 41 47 properties, 129 ratio, 14, 57, 196, 204, 290, 333 335, 339 343, 1001 1002 recommendations, 66 73 reduction factor, 290, 291f rubber, 213 strategy of Nunoa Capital Building, 893 utilization, 736 in tall buildings, 55 73 damping modeling, 55 57 earthquake-excited motion, 85 90 intrinsic damping, 57 73 wind-excited motion, 78 85 technology, 102 103 tensioners, 35 trends, 932 value, 68 Damping system(s), 8, 14, 77, 313, 438, 468, 471 472, 733. See also Buildings over 200 m in United States (Database 2); Testing,

inspection, and maintenance; Worldwide buildings over 250 m (Database 1) additional damping providing by, 737 design, 109, 321 determination of TLCD design parameters, 339 344 harmonic loading and inherent damping, 341t parameters, 294 295 preliminary design, 419 determination of design parameters of TLD, 334 339 drag coefficient values for slat screens, 337f mass ratio, 332 TMD design parameters determination, 333 334 trends, 945 947 for damping system category, 937 941, 947 for damping system types, 941 943, 948 white-noise loading and inherent damping, 342t Damptech, 194 Database source, 929 930 DBE. See Design Basis Earthquake (DBE) D/C ratios. See Demand/capacity ratios (D/C ratios) DE. See Design Earthquake (DE) Dead load, 17 Debonding agents, 191 Deep-water TLD, 201 202 Deficiency-based procedure. See Tier 2 procedure Deflection amplification factor, 284 Deformation, 15 deformation-controlled action, 12 demand, 347 parameters, 90 response factor, 44 and phase angle as function of frequency ratio, 45f Degree of Freedom (DOF), 47, 47f, 371 372 Deicon, 206, 232 Demand/capacity ratios (D/C ratios), 601 602 Dennis Lau & Ng Chun Man Architects and Engineers (DLN), 801 Design Basis Earthquake (DBE), 15, 899 Design codes improvement, 924 925 Design displacement, 15 Design Earthquake (DE), 268 269, 601, 707 floor deflection in modal and total forms, 294 295 roof displacement in modal forms, 294 story drift in modal and total forms, 295 velocity in modal and total forms, 295 Design Engineer, 15 Design force in structural elements, 294

1076 Index

Design lateral force in modal form, 294 Design procedures for tall buildings, 235 236 active, semiactive, and hybrid systems, 365 400 codes and design tools, 236 264 codes and guidelines, 236 247 practical design aspects, 248 251 structural analyses, 251 264 dynamic modification devices strategy optimization, 429 436 isolation systems, 347 365 passive damping systems, 264 347 retrofit of existing buildings, 400 429 Design response spectrum, 270 272, 272f Design review of damping system, 469 Design spectral response acceleration parameters, 270 Design strength, 15 Design wind load cases, 281, 282f Deterministic Risk-Targeted Maximum Considered Earthquake Ground Motions, 273 Development phase, 899 Device control laws, 388 389, 388f Device testing tools, 479 482, 485 486 for damper attached with building structure, 481 dampers tested from building structure, 480 481 for hybrid systems, 481 482 shaking table tests, 485 substructure hybrid tests, 486 wind tunnel test, 485 DFMD. See Dual function metallic damper (DFMD) Diagonal brace dampers, 111 115 Diagonal bracing, 111 115, 187 188 Diagonal dampers, 111 115, 441 442, 443f in diagonal-braced frames, 115f Diagrid frames, 461 systems, 462 Diagrid fac¸ade, 460 462 interaction of facade and dampers in complexshaped building, 461f topmost stories of 72-story building indicating, 463f 20-story base-isolated building in Tokyo, 463f Diaphragm, 15 Direct feedback, 376 Direct-integration analysis, 256 258 Directional wind procedure, 275 281, 276f Discrete systems, 77 78 Displacement, 15 activated supplemental damping system, 254 dependent devices, 245 246

displacement-dependent dampers, 108 109, 141 142, 478, 504, 516 dampers and elastic-perfectly-plastic hysteretic loop, 143f damping device, 15 devices, 479 force, 124 125 hysteretic models of displacement-dependent devices, 142 144 materials, 145 151 metallic yielding devices, friction, 141 195 properties estimation, 308 312 special connection devices, 169 183 special structural members, 184 192 standard devices, 152 169 type of devices and manufacturers, 192 195 feedback, 381 gain, 379 reduction ratio, 240 242 restraint system, 15 Dissipative devices, 239 240 cost implication of buildings with, 101 102 Dissipative elements, 459, 460f Distributed damped systems, 509 Distributed dampers, 478 479. See also Isolated dampers analysis procedure selection, 288 298, 289t available design procedures, 267t building and site categorization, 268 284 building fundamental properties and preliminary structural analyses, 284 288 classification of tall building structural systems, 285f damped structural model construction and perform structural analyses, 312 317 damper type, configuration, and distribution, 298 302 damping preliminary design, 302 312 fully linear distribution of damping along tall building height, 302f lateral force-resisting system selection, 284 load combinations, 283 284 load-resisting systems and design coefficients, 286t quality control, maintenance, and inspection requirements, 321 response acceptability checking, 317 321 risk category and occupancy importance factor, 268 site spectral response acceleration, response spectrum, and time histories, 268 275 step-by-step procedure for, 265 321, 266f total target damping selection, 298 wind demand, 275 283

Index 1077

Distributed damping systems, 109, 418, 439 443, 461, 472 482, 494 495, 516, 534 733. See also Mass damping systems Allianz Tower, Milan, Italy, 716 733 approaches, 109 195 displacement-dependent metallic yielding devices, friction, 141 195 geometrical amplification factor of damper displacement, 114f inclined viscous damper with vertical and horizontal deformations, 115f portal scheme with diagonal viscous damper, 115f San Francisco 181 Fremont Street mega brace frame, 121f velocity-dependent viscous dampers, VE dampers, 126 141 Atushi Building, Xin Jiang, China, 668 677 Beijing Yintai Center, Beijing, China, 577 592 Columbia Tower, Seattle, USA, 534 547 Connor Tower, Manila, Philippines, 701 716 Costums Residential, Auckland, New Zealand, 678 701 damper configurations, 439 device testing tools, 479 482 for damper attached with building structure, 481 dampers tested from building structure, 480 481 for hybrid systems, 481 482 devices, 473 expose or conceal, 441 443 181 Fremont Street, San Francisco, California, USA, 658 668 geometrical configurations and advantages/ disadvantages, 112t Pangu Plaza, Beijing, China, 567 577 San Diego Central Courthouse, San Diego, California, USA, 592 615 St Francis Shangri La Place, Manila, Philippines, 556 566 steel dampers, 492 testing based on Chinese standard, 478 based on European standard, 478 479 based on Japanese standard, 474 478 based on US standard, 473 474 Tianjin International Trade Center, Tianjin, China, 626 637 Two Union Square, Seattle, USA, 547 556 VE dampers, 491 492 vertical and horizontal distribution, 439 441 viscous dampers, 491

Wuhan Poly Cultural Plaza, Wuhan, China, 615 626 454 Yonge, Toronto, On, Canada, 637 658 Distributed systems, 77 Distributed TMDs, 458 459 Distributed viscous wall dampers in SUT Building, 441f Distributed-type dampers, 429, 938 941 test requirements for, 475t DLN. See Dennis Lau & Ng Chun Man Architects and Engineers (DLN) DoCoMo building, 124 DOF. See Degree of Freedom (DOF) Dorman Long Technology, 133 134 Double knee brace systems, 164 Double pendulum approach, 446 Double round-hole metallic damper, 156 Double skin fac¸ade (DSF), 458 459 dissipative elements connecting wall systems and, 460f energy dissipating systems in, 459f interaction of active damping system (damper plus actuator) and, 460f Double X-shaped damper, 156 Double-acting device, 180 Double-layer shear damper, 29 Double-tee BRB systems, 191, 193f Double-tube BRB systems, 191, 193f Double-vane system, 134 Double/Triple/Quintuple friction pendulum, 361 “Double” isolators, 214 pendulum system principles and its force displacement behavior, 215f Douce Hydro, 134 Drift, 20 criteria, 319, 346 ratio, 20 DSF. See Double skin fac¸ade (DSF) DTU. See Technical University of Denmark (DTU) Dual function metallic damper (DFMD), 152, 155 156, 159f Dual system, 16, 566, 658 Dual-use TMD, 908 909, 918 Ductility, 15, 145 demand determination, 296 Duhamel’s integral, 46 “Dunnage” beams, 751 753 Dynamic behavior, 79 Dynamic element, 20 Dynamic force, 30 31 Dynamic Isolation System, 134, 216, 898

1078 Index

Dynamic linear analysis, 238 procedure, 419 420 Dynamic models and simulation techniques, 453 Dynamic modification devices, 923 active, semiactive, and hybrid systems, 217 232, 222t base isolation, 207 217 comparison of dampers in tall buildings, 232 passive damping systems, 108 206 strategy optimization, 429 436 algorithm-based optimization procedures, 430 432 nonalgorithm-based optimization procedures, 433 436 type of dampers, 429 in tall buildings, 28 30 typologies of damper devices, 108t Dynamic modification systems implementing results of technological progress, 925 928 development of existing devices and new hybrid systems, 925 926 new technologies, 926 928 improved modeling of structural behavior, 922 925 enhanced computing power, 923 924 improve behavior model of materials and devices, 923 improve understanding of actions on structures, 922 923 improved design codes, 924 925 Dynamic response factors, 44 45 Dynamic seals, 131 132 Dynamic tests, 15, 484 Dynamic wind-induced excitation, 78

E E-PAD, 194 Earthquake Engineering Research Institute (EERI), 508 Earthquake(s), 454 actions, 66 damping, 66 earthquake-resisting systems, 185 excitation, 379 observed response under, 508 intensity levels, 246 EBB. See Euler Bernoulli beam (EBB) Eccentrically braced frame (EBF). See Concentrically braced frame (CBF) Eccentricity, influence of, 281 ECD systems. See Eddy current damper systems (ECD systems)

Eddy current, 927 928 damping force, 872 Eddy current damper systems (ECD systems), 927 928, 927f EDP. See Engineering demand parameter (EDP) EDR. See Energy dissipating restraint (EDR) EDWG. See Energy Dissipation Working Group (EDWG) EERI. See Earthquake Engineering Research Institute (EERI) Effective damping, 13, 798 Effective ductility demands, 292 Effective fundamental mode period, 292 Effective stiffness, 16 for RC components, 262 263 Effective water mass ratio, 833 Efficiency index, 339 340 Eigen-Vectors, 259 Elastic buckling strength, 185 deformation, 15 seismic demand, 353 stiffness, 358 unloading, 146 Elastic-perfectly-plastic behavior, 156 160 Elastic-perfectly-plastic model, 142 displacement-dependent dampers and, 143f Elastomeric bearings, 212 213, 236 237 isolators, 212 spring damper, 180 Electrolytic tough pitch copper (ETP copper), 148 Electrorheological damper (ER damper), 227 228, 518 Elevators, 450 454 designers, 453 installations, 451 ELF. See Equivalent lateral force (ELF) Elliptical diagrid systems, 461 ELM. See Equivalent linearization method (ELM) Empirical stiffness modifiers, 254 EN 15129 Standard, 131, 478 479, 504 Encasing tube, 191 Enclosure classification, 279 Energy dissipating restraint (EDR), 178, 181f Energy dissipation, 15, 108 109, 148 capacity, 160 device, 15 buildings with, 245 selection, 298 299 mechanism, 164 systems, 15, 243 Energy Dissipation Working Group (EDWG), 246 247 Engineering demand parameter (EDP), 101

Index 1079

Enhanced computing power, 923 924 “Enhanced” seismic performance objective, 610 611 Environmental and economic considerations of building life cycle, 90 102 cost implication of buildings with dissipation devices, 101 102 LCA, 93 101 Equations of motion, 41 42, 47 48 Equivalent damping coefficient for sinusoidal excitation, 338 formulation comparison, 974f parameters, 693 Equivalent lateral force (ELF), 288 289 Equivalent linearization method (ELM), 1012 Equivalent mechanical models, 203 204 Equivalent viscous damping, 13, 53 54, 980f ER damper. See Electrorheological damper (ER damper) ETP copper. See Electrolytic tough pitch copper (ETP copper) Euler Bernoulli beam (EBB), 260 261 European code, 236 238 Eurocode 1, 66 68, 67t Eurocode 8, 347 348, 367, 389 European Conformity, 728 730 European standards, 472 testing, 478 479 Evaluation of damping constant for Fluid damper, 976f of existing building, 410 412 experimental evaluation for isolated dampers, 346 347 procedures based on ASCE 41 13, 402 403 Excitation, 31 “Expected strength”, 254 255, 415 Exposed dampers, 439 Exposed distributed damping systems, 441 443 externally exposed distributed dampers, 442f internally exposed distributed dampers, 442f Exposure category, 276 Extension stroke, 128 129 Extensive temporary instrumentation system, 900 Exterior Mega Frame, 865 External pressure coefficient, 281 EZFRISK computer software, 274

F Fac¸ade, 457 464 diagrid fac¸ade, 460 462 DSF, 458 459 mega brace dampers fac¸ade, 463 464 Factory production control test (European Standard). See Production test

Factory test, 16 FAM. See Fluid-based Aerodynamic Modification (FAM) Fast Fourier transform (FFT), 816 Fast nonlinear analysis (FNA), 253, 256 259 Fatigue of dampers, 501 505 base isolation systems, 505 displacement-dependent dampers, 504 isolated damping systems, 504 viscoelastic dampers, 503 viscous dampers, 502 503 endurance limit, 16 FDDR. See Free decay damped response (FDDR) FEA. See Finite element analysis (FEA) Federal Emergency Management Agency (FEMA), 636 FEMA 450 requirements, 246 247 FEMA 751, 360 361 FEMA P-1050 1, 347 348 FEMA-178 standard, 400 401 Feedback control, 218 Feedforward control, 218 FEMA. See Federal Emergency Management Agency (FEMA) FFT. See Fast Fourier transform (FFT) Fiber Technology Corporation, 206 54-story building, 509, 509f Filler material, 191 Finite difference method, 202 203 Finite element analysis (FEA), 252 software packages, 923 924 Finite elements, 3 4 method, 202 203 Finite volume method, 202 203 FIP Industriale, 134, 141, 194 195, 206, 216 217 First Hills Libadashi building, 449 “Fisher” information, 62 Flag-shaped behavior, 176 Flat slab, 16 Flat sliding bearings, 214, 214f Flexible metal bellows, 132 133 Flexural deformations, 111 115 Floating slab track, 34, 36f Floor spectral acceleration, 98 100 Fluid dampers, 126 127, 128f, 132, 976f Fluid models, 202 203 Fluid structure coupling (FSC), 928 Fluid viscous dampers (FVDs), 14, 29, 126 136, 572, 627, 926, 971 975 additional considerations, 131 133 design description, 128 129 functional description, 129 130 hysteresis loop for, 972f specification requirements, 131

1080 Index

Fluid viscous dampers (FVDs) (Continued) types of devices and manufacturers, 133 136 temperature dependency, 975 Fluid-based Aerodynamic Modification (FAM), 106 FNA. See Fast nonlinear analysis (FNA) Force coefficient, 317, 318t Force reduction ratio, 240 242 Force-controlled action, 12 Force-resisting system selection, 326 Forced vibration properties, 43 47, 53 tests, 16, 499 500 Force displacement, 142 143 hysteretic response, 137 138 relationship, 976 977 Four-DOF model, 376 Fourier transform of excitation, 370, 371f FPB. See Friction pendulum bearings (FPB) FPC, 194, 216 Frame, 16 braced, 16 braced moment, 540 Chevron, 116 diagrid, 461 Exterior Mega Frame, 865 frame-action, 58 60, 60f mega, 958 moment, 16 steel, 735 systems, 958 toggle, 116 117 Free decay damped response (FDDR), 61 62 Free sliding bearings, 214 Free vibration properties, 42 43, 48 52 tests, 16, 483, 500 181 Fremont Street, San Francisco, California, USA building overview, 659f comparison different design strategies damping solution consideration, 667 structural options consideration, 667 damper and brace installation, 665f damper main characteristics, 665t damping overview architectural integration strategy, 666 damping strategy, 662 664 damping type, 664 665 elevator and other mechanical devices integration strategy, 667 experimental tests, 667 maintenance strategy, 667

MEP integration strategy, 667 monitoring system, 667 structural and damping design, 665 666 introduction/history, 658 megabrace damper 3D representation, 664f megabrace damper configuration, 664f project data, 658 lesson learned and recommendations, 668 structural elevations, 662f structural floor plan, 660f structural system, 658 662 additional damping provided by damping system, 661 building code, 661 building cost vs. damping cost, 661 building fundamental periods, 660 damping strategy utilization, 660 design forces, 662 expected performance, 662 peer-reviewed project, 661 3d structural model, 663f Fremont tower, 464, 464f French code, 72 Frequency response function (FRF), 61 62 Frequency-dependent ISO criteria, 791 792 Freyssinet, 134 FRF. See Frequency response function (FRF) Friction coefficient, 214 cyclic behavior, 981 damping. See Coulomb damping force, 170 and hysteretic dampers modeling, 314 315 pad disks, 172 173 tests, 606 608 Friction dampers, 14, 109, 128, 169 176, 309 310, 522, 981 983 behavior, 170 pall friction damper, 170f pall system, 171 172 rotational friction system, 172 176 slotted-bolted connections, 171 sumitomo friction devices, 171 system, 983 Friction pendulum bearings (FPB), 214, 214f, 236 237, 462, 823 Frictionless hermetic dampers, 120 121 FSC. See Fluid structure coupling (FSC) Full-scale VCD test, 138 in uniaxial configuration subject to simulated earthquake, 139f Functional description, 129 130, 137 138 Fundamental period, 16

Index 1081

Fuzzy logic control, 231 FVDs. See Fluid viscous dampers (FVDs)

G GA. See Genetic algorithm (GA) GB 50011 standard, 243, 245 246, 471 Generalized Maxwell Model (GMM), 711 Genetic algorithm (GA), 228, 325 326, 430 432 Geometrical configuration, 439 Geometrical magnification factor, 308 Gerb, 206, 232 GERB Schwingungsisolierungen Gmbh & Co, 134 Germany DIN1055 Teil 4 code (DIN 2005), 68 Giant truss, 449 450 Glass square tower, 735 Global energy consumption, 93 Global Navigation Satellite System, 916 Global property modification factor, 264 GMM. See Generalized Maxwell Model (GMM) Gradient-based algorithms, 430, 431f Gravity system, 16 Grip supplies energy, 926 Ground motions, 274 275 intensity, 351 selection and scaling of ground motion records, 274 275 Gusset plates, 191 Gust effect factor, 276 279, 278t Gust loading. See Seismic loading

H Hairpin shape steel “springs”, 742 Handover procedure, 526 Harbin Institute of Technology (HIT), 580 Hardware, 438 Harmonic, 30 31 excitations, 381 382, 397, 985 SDOF response, 992f TMDs SDOF simplified procedure for, 993 995 motion, 991 vibration tests, 483 Harvesting system, 927 928 HDRB. See High damping rubber bearing (HDRB) HDS. See Honeycomb damper (HDS) Head loss coefficient, 339 340 Heat transfer process, 133 Hermetic dampers, 135 136 Heuristic algorithms, 430, 432 High damping rubber, 139, 140f High damping rubber bearing (HDRB), 139, 349 High-hardness rubber, 136 137 High-performance-damping systems, 921 922

High-rise buildings, 57 primary source of data for measurements of mode intrinsic damping, 59f retrofit strategies for 24-story building in Osaka with oil dampers, 425 for 27-story building in Osaka TMD, 428 429 for 30-story building in Osaka with viscous dampers, 425 428 for 35-story building in San Francisco with viscous dampers, 422 425 Highcliff Apartments, Hong Kong, USA, 801 building accelerometers, 807f comparison different design strategies cost benefit analysis, 808 damping solution consideration, 808 structural options consideration, 807 damping overview architectural integration strategy, 806 elevator and mechanical devices integration strategy, 806 experimental tests, 806 MEP integration strategy, 806 monitoring system, 806 807 strategy, 805 structural and damping design, 805 806 type, 805 floor plan at mechanical penthouse showing damper tanks, 804f fundamental frequencies and mode shapes, 804f introduction/history, 801 802 lesson learned and recommendations, 808 project data, 801 structural system, 803 805 additional damping provided by damping system, 804 building code, 804 building cost vs. damping cost, 804 building fundamental periods, 803 damping strategy utilization, 803 peer-reviewed project, 805 TLD internal view of dissipating screens, 806f schematic illustration, 805f typical tower floor, 803f wave sensor, 807f on roof, 808f Higher-modes, 923 Higher-order DOF systems, 375 376 HIT. See Harbin Institute of Technology (HIT) HMD. See Hybrid mass damper (HMD) Hoistway equipment, 453 Honeycomb damper (HDS), 155, 156f, 157f

1082 Index

Hong Kong code of practice on wind effects, 68 Horizontal distribution, 329 330, 439 441 Horizontal friction coefficient, 740 742 Horizontal positions, 439 Hourglass-shaped strip dampers (HSDs), 160 H-Safe, 133 HSDs. See Hourglass-shaped strip dampers (HSDs) Human-powered vibration test, 16 Hummingbird Kinetics, 206 Hyatt Park Tower, Chicago, USA, 789f, 802f comparison different design strategies, 799 800 cost benefit analysis, 800 damping solution consideration, 800 structural options consideration, 800 damping overview, 792 799 architectural integration strategy, 799 elevator and mechanical devices integration strategy, 799 experimental tests, 799 maintenance strategy, 799 MEP integration strategy, 799 monitoring system, 799 strategy, 793 794 structural and damping design, 795 799 type, 794 795 introduction/history, 788 789 lesson learned and recommendations, 800 801 project data, 788 structural system, 790 792 additional damping provided by damping system, 790 building code, 791 building cost vs. damping cost, 790 791 building fundamental periods, 790 damping strategy utilization, 790 expected performance, 791 792 peer-reviewed project, 791 TMD acceleration performance, 792f characteristics, 793t scheme, 794f wind tunnel test, 796f Hybrid control systems, 219 220, 220f, 225 Hybrid dampers, 518 519 Hybrid damping systems, 492 493, 497t Hybrid mass damper (HMD), 220, 225, 225f, 370 371, 512 513, 908 Hybrid system, 517 518, 534 active mass damper, 510 actuators, 223 224 ATMD, 222 223, 229 230, 510 512, 511f case studies, 904 919

Thyssenkrup Test Tower, Rottweil, Germany, 904 919 control strategies, 230 231 dampers manufactures, 232 future directions, 231 232 HMD, 512 513 interaction of active damping system and DSF systems, 460f literature review, 365 367 semiactive, hybrid dampers manufactures, 232 step-by-step procedure, 367 400, 368f bandwidth specification of active and semiactive systems, 370 build simplified and reduced model, 372 376 building and site categorization, 369 building fundamental properties and preliminary structural analyses, 369 370 damper type, configuration, and distribution of sensors and actuators, 370 372 design of controller system, 372 376 selecting structural system, 369 selection of performance objective, 370 verification analyses and quality control, 394 400 Hybrid systems, 180 182, 182f, 214, 217 232, 499 500 ATMD systems, 229 230 control strategies, 230 231 damper manufactures, 232 future directions, 231 232 HMD, 225 semiactive control of base isolation systems, 225 testing tools, 481 482 Hybrid test. See Sub-structures online test Hybrid testing, 18, 479 480, 482f Hydraulic actuator, hybrid testing with, 486 Hydraulic axial dampers, 78 Hydraulic cylinder, 132 Hydraulic dampers, 129 design, 795 Hydraulic systems, 132 Hydro-pulse cylinders, 223 Hysteretic/hysteresis, 16 behavior, 58 60, 142 143, 148 of steel, 146 curve, 161 163 dampers, 130, 310 312, 473, 978 981 fluid, 129f SDOF response with, 980f damping, 13, 55 model, 54 ratios determination, 296

Index 1083

devices, 212 213 energy dissipation, 979 models of displacement-dependent devices, 142 144

I IBC. See International Building Code (IBC) IFDL system. See Inertia-force-driven loading system (IFDL system) Immediate occupancy, 18, 925 IMPACT 2002 1 framework, 97 98, 97f Importance factor, 16 Impulse response function (IRF), 61 62 The Independent, Austin, Texas, USA, 877f, 878f, 879f, 887t comparison different design strategies cost benefit analysis, 887 damping solution consideration, 887 structural options consideration, 887 damping overview, 882 886 architectural integration strategy, 886 development process, 886 elevator and mechanical devices integration strategy, 886 experimental tests, 886 maintenance strategy, 886 MEP integration strategy, 886 monitoring system, 886 strategy, 883 structural and damping design, 885 type, 884 decay plot showing damper performance, 885f fundamental frequencies and mode shapes, 880f introduction/history, 877 lesson learned and recommendations, 887 888 design innovative solutions, 888 difficulties in design, 888 improvements, 888 project data, 876 877 structural axis convection, 882f structural system, 877 882 additional damping provided by damping system, 880 building code, 880 building cost vs. damping cost, 880 building fundamental periods, 879 damping strategy utilization, 879 880 design forces, 881 expected performance, 881 882 peer-reviewed project, 881 wind tunnel test, 881f Indirectly connected dampers, 439

Industrial application of damping devices, 30 37 industrial machine structures, vibration and seismic control of, 32 33 bearing supported fans in air conditioning system and power plant, 32f elastic support of diesel engine principle, 33f turbine generator machine buildings, 33f vibration isolation of transformer and with isolators, 34f vibration control, 30 31 of large industrial chimney stacks, 34 37 ring-type tuned mass damper, 37f vibration isolation of buildings and rail structures, 34 Inelastic deformation, 15 Inertance, 928 Inerter, 928, 928f Inertia-force-driven loading system (IFDL system), 482, 483f Inertial forces, 30 31 Infilled shear wall. See Shear panel damper (SPD) Inherent damping. See Intrinsic damping Inspection, periodic testing, and maintenance program, 469 470 Installation admissible error, 525 Inter-story drift, 16 proportional distribution, 434 ratio, 98 100 Intermediate-water depth TLD, 201 202 Internal pressure coefficient, 279, 280t International Building Code (IBC), 706 707 International Organization for Standardization (ISO), 84, 791 792, 832 833, 843, 850 851 Intrinsic damping, 13, 55, 57 73 code limitations, 66 73 estimation, 61 62 prediction, 62 66 sources, 57 60 Inversely linear damping distribution, 310 Inversely linear distribution, 306 Inverted V-Bracing, 176, 187 188 IRF. See Impulse response function (IRF) ISD-111H material, 503 ISO. See International Organization for Standardization (ISO) ISO 4354 code, 68 intrinsic damping ratio recommendation, 69t ISO/CD 3010 code, 68 Isolated damped/damping systems, 14, 492, 504, 510, 516 517, 938 941 active, semiactive and hybrid systems, 499 500 device testing tools, 485 486

1084 Index

Isolated damped/damping systems (Continued) shaking table tests, 485 substructure hybrid tests, 486 wind tunnel test, 485 passive, active, semiactive and hybrid damping systems, 482 486 MR dampers, 484 TLDs/TLCDs, 484 TMDs, 483 484 TLCDs, 499 TMDs, 495 498 Isolated dampers, 938, 941 943, 950 953. See also Distributed dampers analyses procedure selection, 327 building and site categorization, 326 building fundamental properties and preliminary structural analyses, 326 327 damper type, configuration, and distribution, 328 331 damping system preliminary design, 331 344 design considerations, 321 326 available procedures, 324 326, 325t TLCD, 323 324 TLD, 322 323 TMD, 321 322 force-resisting system selection, 326 quality assurance and experimental evaluation, 346 347 response acceptability checking, 346 step-by-step procedure, 321 347 total target damping selection, 327 328 update building model and perform analyses, 344 346 Isolation at base, 447 above cantilever floor columns, 447, 448f device, 17, 236 237 above first-story walls, 448, 449f interface, 16 layer, 208, 449 450 plane, 447, 447f story, 449 Isolation systems, 17, 209, 237, 243, 347 365, 468, 895 898. See also Passive damping systems modeling, 360 363 primary design, 357 359 step-by-step procedure for base isolation, 347 365 type and distribution, 351 353 determine type, 351 353 location and distribution, 353 ISOSISM, 216 Isosism FD, 134

Italian Code, 72, 238, 347 348 ITT Endine, 134

J J2 Building of Tokyo Institute of Technology, 210 Japan, damped buildings in, 935 936 Japan Society of Seismic Isolation (JSSI), 238 239, 239f, 241f, 508 Japanese Building Disaster Prevention Association (TJBDPA), 401 Japanese Building Standard Law (BSL), 238 239 Japanese code, 73, 238 242, 347 348, 470 471 Japanese standard, 504 code, 501 502 testing, 474 478 steel dampers, 477 478 VE dampers, 477 viscous dampers, 476 JCC. See Judicial Council of California (JCC) Jeary’s theoretical damping model, 62f JG/T209 2012 standard, 471 JGJ297 2013 standard, 471 Jinan Baidungs Machinery Equipment Co., Ltd, 134 John Hancock Tower, Boston, USA, 198, 748f, 749f, 750f, 751f, 752f comparison different design strategies cost benefit analysis, 757 damping solution consideration, 757 structural options consideration, 757 damping overview architectural integration strategy, 755 development process, 756 elevator and mechanical devices integration strategy, 755 756 experimental tests, 756 maintenance strategy, 756 MEP integration strategy, 755 monitoring system, 756 strategy, 750 751 structural and damping design, 753 755 type, 751 753 introduction/history, 746 748 lesson learned and recommendations, 757 project data, 745 746 structural system, 748 750 additional damping provided by damping system, 749 building code, 749 building cost vs. damping, 749 building fundamental periods, 748 damping strategy utilization, 748 749 design forces, 750

Index 1085

expected performance, 750 peer-reviewed project, 749 JSSI. See Japan Society of Seismic Isolation (JSSI) Judicial Council of California (JCC), 593 “enhanced” seismic performance objectives, 600 601

K Kawakin Core-Tech Co., Ltd., 195 KBF. See Knee Bracing System (KBF) Kelvin model, 314 solid model, 137, 137f Kinetica, 141 Knee Bracing System (KBF), 164 166 configurations, 165f Knee element, 164 165

L L-tower, Toronto, Canada, 829f bidirectional TLSD, 834f comparison different design strategies cost benefit analysis, 839 damping solution consideration, 839 structural options consideration, 839 coordination axis, 831f damping overview, 833 839 architectural integration strategy, 836 837 development process, 839 elevator and mechanical devices integration strategy, 837 experimental tests, 837 838 maintenance strategy, 838 839 MEP integration strategy, 837 monitoring system, 838 strategy, 833 structural and damping design, 834 836 type, 833 detail damper drawing, 836f elevations, 830f fundamental frequencies and mode shapes, 831f introduction/history, 828 lesson learned and recommendations, 839 pendulum-type dynamic test rig, 838f project data, 828 structural system, 828 833 additional damping provided by damping system, 832 building code, 832 building cost vs. damping cost, 832 building fundamental periods, 828 831 damping strategy utilization, 832 design forces, 832

expected performance, 832 833 peer-reviewed project, 832 typical floor plan, 830f LA City Hall, 209 Large lateral displacements, 454, 455f Large mass, 444 445 Large-scale testing approach, 481 LATBSDC. See Los Angeles Tall Buildings Structural Design Council (LATBSDC) Lateral force-resisting system parameters of, 292 294 selection, 284 Lateral load, 17 resisting system, 252 LCA. See Life cycle assessment (LCA) LCVA, 205 LDRBs. See Low-damping rubber bearings (LDRBs) Lead, 148 damper hysteresis loop with insignificant aging effects, 148f dampers, 14, 167f, 168 169, 505, 523 damage to, 523f extrusion dampers, 167 Lead extrusion devices (LEDs), 167 169, 168f Lead rubber bearings (LRBs), 210, 213, 213f, 349, 351 352, 505, 823 LEDs. See Lead extrusion devices (LEDs) LeMessurier Associates, 744 Lever damper, 119, 119f 601 Lexington. See Citicorp Center Life cycle assessment (LCA), 17, 90 101 damaged-oriented, 97 101 Life safety (LS), 18, 720, 925 Lifts. See Elevators Lightly tensioned traveling cables, 451 453 Limit state, 17 Linear analysis, 898 899 methods, 252 255 procedures, 419 420 Linear approximation method, 389 Linear dampers, 129, 566 Linear dynamic analysis methods, 253 Linear fluid models, 202 203 Linear force velocity relationship, 130, 130f Linear hydraulic dampers, 798 Linear methods for supplemental damping systems, 253 255 Linear motors. See Hydro-pulse cylinders Linear quadratic Gaussian (LQG), 230 231 Linear quadratic regulator (LQR), 230 231, 366 algorithm, 383 384 approach, 369

1086 Index

Linear static analysis, 253 Linear time-history analysis methods, 253 Linear viscous dampers (LVDs), 130 131 Linear viscous damping system, 973 Linked stabilizing/destabilizing forces principle, 814 815 Liquid column length, 343 344 of each group, 344 damper, 883, 883f, 884f depth ratio, 335 spring-damper component, 178 pressurized fluid restoring device including, 180f spring-hardening, 322 323 turbulence, 444 445 velocity standard deviation, 1009 Live load, 17 Load combinations, 283 284, 316 317 factor, 4, 17 and forces, 17 path, 17 resistance, 748 Local buckling, 191 Logarithmic decrement, 72 Los Angeles Tall Buildings Structural Design Council (LATBSDC), 71 72, 706 707 Loss factor, 978 Low-damping NRBs, 212 213 Low-damping rubber bearings (LDRBs), 210 Low-pass filtering, 376 Low-stiffness connectors, 458 Low-yield-strength steel (LYS 100), 504 Lower and upper specification tolerance, 264 Lower bound period, 274 LQG. See Linear quadratic Gaussian (LQG) LQR. See Linear quadratic regulator (LQR) LRBs. See Lead rubber bearings (LRBs) LS. See Life safety (LS) LVDs. See Linear viscous dampers (LVDs) LYS 100. See Low-yield-strength steel (LYS 100)

M Machine-Room-Less (MRL), 454, 454f Magnetic circuit design, 387 Magnetic field test, 484 Magnetorheological dampers (MR dampers), 228, 397 399, 484 design, 384 387 shear-valve MR damper, 385f fluids, 227 228 mechanical model, 229f

responses, 229f shear mode, 228f Magnusson Klemencic Associates (MKA), 539, 547, 644, 654, 707 Malaysia Rubber Producers’ Research Association (MRPRA), 213 Martensite, 151 Mass, 1005 block-building interface, 743 dampers, 14, 457 mass-type dampers, 429 ratio, 196, 203, 205 206, 332 of TMDs, 333 334 Mass damping systems, 419, 444 446, 446f, 462, 534. See also Distributed damping systems approaches, 195 206 mass damping approaches simplified schemes, 197f TLDs, 200 206 TMD, 195 200 TMD/TLD/TLCD manufactures, 206 Bloomberg Tower, New York city, USA, 808 818 Citicorp Building, New York, USA, 734 745 Comcast Center, Philadelphia, USA, 783 788 Highcliff Apartments, Hong Kong, USA, 801 808 Hyatt Park Tower, Chicago, USA, 788 801 The Independent, Austin, Texas, USA, 876 888 John Hancock Tower, Boston, USA, 745 757 L-tower, Toronto, Canada, 828 839 One Bloor Street East, Toronto, Canada, 847 854 One Rincon Hill (South Tower) San Francisco, California, USA, 774 783 Raffles city, Chongqing, China, 818 827 Shanghai Tower, Shanghai, China, 863 876 Taipei 101, Taipei, Taiwan, 758 774 14 York Street, Toronto, Canada, 840 847 1151 West Georgia, Vancouver, Canada, 854 863 Mass proportional damping distribution (MPD distribution), 434, 434f Material(s), 142, 145 151 aluminum, 149 151 behavior, 138 copper, 148 improving behavior model of materials and devices, 923 lead, 148 material-based dissipation systems, 8 9 SMA, 151 steel, 146 148

Index 1087

strength, 254 255 technologies, 926 testing, 474 477 MATHEMATICA software, 367 369 MATLAB software, 367 369 Matrix-form equation, 53 Maurer, 134 135, 206, 217, 232 Maurer hydraulic damper (MHD), 134 135 Maximum considered earthquake (MCE), 15, 17, 69 71, 473 474, 601, 899 Maximum displacement, 15 Maximum inter-story velocity, 303 305 Maximum rated force, 131 Maxwell model, 314, 632 Maxwell spring effect, 559 MCE. See Maximum considered earthquake (MCE) MCFT. See Modified Compression Field Theory (MCFT) MDOF systems. See Multiple Degrees of Freedom systems (MDOF systems) Mean recurrence interval. See Return period Mean-square responses determination, 396 397 Mechanical, electrical, and plumbing systems (MEP systems), 32, 77, 98, 437 438 Mechanical components, 456 Mechanical devices, 178 Mechanical floors, 456 457 Mechanical systems, 455 457 basic considerations, 455 456 mechanical floors, 456 457 Mega brace dampers fac¸ade, 463 464 Mega frame, 958 Mega tall building, 949 950 Megastructural system, 762 Megastruss, 208 Melted dynamic seal. See Softened dynamic seal MEP systems. See Mechanical, electrical, and plumbing systems (MEP systems) Metallic Materials Properties Development and Standardization (MMPDS), 502 Metallic/metal bellow seals, 129, 135 136 to metal, 171 system development, 29 to teflon, 171 yield damper, 471 yielding damper, 14 MHD. See Maurer hydraulic damper (MHD) Middle-story isolation, 448 449, 455f Mixed-structure, 953 Mixed-Use tall building, 958 961 MKA. See Magnusson Klemencic Associates (MKA)

MMPDS. See Metallic Materials Properties Development and Standardization (MMPDS) Modal damping, 13, 56 Modal effective damping for design and maximum ground motions, 296 297 Modal effective seismic weight, 292 293 Modal filtering, 376 Modal seismic base shear, 292 293 response coefficients, 293 Modal-based equations, 995 Mode coupling, 31 for spring-supported systems, 31f Gakuen Cocoon Tower, 461, 462f participation factor, 51 shape(s), 48 of building, 51 Model/modeling points, 373 reduction, 372 373 relationship between controllability/ observability and locations, 374f SDOF system, 373 375 Modified Compression Field Theory (MCFT), 262 Moment frame, 16 Motion(s), 39 activated devices, 254 motion-dependent dampers, 108 109 MPD distribution. See Mass proportional damping distribution (MPD distribution) MPS systems. See Mechanical, electrical, and plumbing systems (MEP systems) MR dampers. See Magnetorheological dampers (MR dampers) MRL. See Machine-Room-Less (MRL) MRPRA. See Malaysia Rubber Producers’ Research Association (MRPRA) MTLCDs. See Multiple TLCDs (MTLCDs) MTLDs. See Multi TLDs (MTLDs) MTMDs. See Multiple tuned mass dampers (MTMDs) MTS System Corporation, 206 Multi TLDs (MTLDs), 330 Multiobjective LQG methodology, 372 Multiple Degrees of Freedom systems (MDOF systems), 17, 41, 47 53 base-isolated system, 1009 equations of motion, 47 48 forced vibration properties, 53 free vibration properties, 48 52 stick frame system, 238 239, 239f TMD, 983

1088 Index

Multiple mass dampers, 329 331 Multiple TLCDs (MTLCDs), 205, 339 344 SDOF structure with, 206f Multiple tuned mass dampers (MTMDs), 329 330, 333 334, 983 optimum MTMD parameters for damped structure modeled as SDOF system, 1003t system, 1002 theory, 1001 1003, 1002f Multistory X-bracing, 187 188

N Nakanoshima Festival Tower, 208, 209f, 449 450, 451f Nanomaterials, 927 NASA’s Marshall Space Flight Centre, 928 National Building Code (NBC), 925 National Building Code of Canada (NBCC), 81 82, 646 National code and guideline recommendations, 506 National Institute of Standards and Technology (NIST), 922 923 Natural rubber bearings (NRBs), 208 209 Natural vibration modes, 48, 49f NBC. See National Building Code (NBC) NBCC. See National Building Code of Canada (NBCC) NEHRP, 263 264 “Nested TMD” solution, 687 Neural network controllers, 231 New hybrid systems, existing device development and, 925 926 New Zealand code, 246 New Zealand Concrete Structure Standard (NZS), 246 New Zealand Department of Scientific and Industrial Research, 210 Newmark’s method, 46 47 Nickel-titanium, 151 Nippon Steel, 191, 195 and Sumikin Engineering, 141 NIST. See National Institute of Standards and Technology (NIST) NLLINK element. See Nonlinear two node link element (NLLINK element) NLRHA. See Nonlinear response history analysis (NLRHA) NLVDs. See Nonlinear viscous dampers (NLVDs) Nominal damping coefficient, determination of, 390 Nominal load, 17

Nominal strength, 17 Nonalgorithm-based optimization procedures, 433 436. See also Algorithm-based optimization procedures added damping ratio map for structure, 435f types of mass proportional damping systems, 434f Nonisolated structure, 208 Noniterative optimal design for brace-damper system, 435 436 Nonlinear analysis, 55 56, 237 238, 255 260, 898 899 behavior, 256, 926 characteristics, 794 dampers, 129, 566 fluid models, 202 203 force-displacement hysteretic behavior of FPB, 823 procedures, 297 298, 420 response-history procedure, 19, 297 static procedure, 17, 297 298 time history analysis, 360 FEA, 262 Nonlinear response history analysis (NLRHA), 564, 605, 610 611, 665 Nonlinear two node link element (NLLINK element), 314 Nonlinear viscous dampers (NLVDs), 130 131 Nonplanar responses, 453 Nonstructural components, 456 NRBs. See Natural rubber bearings (NRBs) NTC, 238 Number of tanks, 335 Numerical and analytical fluid models, 202 203 Nunoa Capital Building, Santiago, Chile, 210, 889f comparison different design strategies cost benefit analysis, 901 damping solution consideration, 901 structural options consideration, 901 damping overview, 895 901 architectural integration strategy, 899 damping strategy, 895 development process, 901 elevator and other mechanical devices integration strategy, 899 900 energy dissipation alternatives, 896f experimental tests, 900 maintenance strategy, 900 901 MEP integration strategy, 899 monitoring system, 900 structural and damping design, 898 899 type, 895 898

Index 1089

flexible for fir extinguishing system, 900f fundamental frequency with and without base isolation, 892f introduction/history, 890 lesson learned and recommendations design innovative solutions, 904 difficulties in design, 901 903 possible improvements, 904 plan layout of isolated towers, 891f project data, 888 890 seismic isolation layout, 893f structural system, 890 895 additional damping, 894 building code, 894 building cost vs. damping cost, 894 building fundamental periods, 890 893 damping strategy, 893 design forces, 894 expected performance, 895 peer-reviewed project, 894 NZS. See New Zealand Concrete Structure Standard (NZS)

O Oasys GSA, 564 OBS criterion. See Optimum brace stiffness criterion (OBS criterion) OCC. See Occupant Comfort (OCC) Occupancy, 17 importance factor, 268 Occupant Comfort (OCC), 80 84, 686 code perception criteria, 81 84 process, 249 250 wind loads, 686, 691 ODS criterion. See Optimum damper size criterion (ODS criterion) Oil dampers, 14, 449 450 retrofit strategies for 24-story building in Osaka with, 425 Oiles viscous wall damper, 135, 217 One Bloor Street East, Toronto, Canada, 847f comparison different design strategies cost benefit analysis, 854 damping solution consideration, 853 structural options consideration, 853 damping overview, 851 853 architectural integration strategy, 852 development process, 853 elevator and mechanical devices integration strategy, 853 experimental tests, 853 maintenance strategy, 853 MEP integration strategy, 852

monitoring system, 853 strategy, 851 structural and damping design, 852 type, 851 852 introduction/history, 847 848 lesson learned and recommendations, 854 project data, 847 848 structural system, 848 851 additional damping provided by damping system, 850 building code, 850 building cost vs. damping cost, 850 building fundamental periods, 848 849 damping strategy utilization, 850 design forces, 850 expected performance, 850 851 peer-reviewed project, 850 three-dimensional scheme of bidirectional sloshing damper, 851f typical elevation, 849f typical floor plan, 848f One Rincon Hill (South Tower) San Francisco, California, USA, 775f, 776f, 777f building accelerometer, 781f comparison different design strategies cost benefit analysis, 782 783 damping solution consideration, 782 structural options consideration, 782 damping overview architectural integration strategy, 781 elevator and mechanical devices integration strategy, 781 experimental tests, 781 MEP integration strategy, 781 monitoring system, 781 strategy, 779 structural and damping design, 780 781 type, 780 fundamental frequencies and mode shapes, 778f introduction/history, 776 measured building response to South Napa Earthquake, 782f project data, 774 775 recommendations design innovative solutions, 783 difficulties in design, 783 improvements, 783 shake table test of TLD at UWO, 780f south tower buckling restrained brace installation, 778f structural system, 776 779 additional damping provided by damping system, 777 building code, 779

1090 Index

One Rincon Hill (South Tower) San Francisco, California, USA (Continued) building cost vs. damping cost, 777 778 building fundamental periods, 776 damping strategy utilization, 777 peer-reviewed project, 779 Online hybrid test, 18, 491 OOME. See Output-only modal estimation (OOME) OP/OTP devices. See Fluid viscous dampers (FVDs) Operational Modal Analysis, 733 Optimal closed-loop control scheme, 365 Optimal displacement feedback control law, 366, 379 381 Optimal KBF shape, 164 Optimal placement with controllability index, 434 Optimal tuning frequency, 333 335, 339 343 Optimality with maximization of supplemental damping ratio, 433 434 of mid-story isolation systems parameter, 436 of TLCD parameters, 436 of TMD tuning ratio, 436 Optimization strategies, 429 Optimum brace stiffness criterion (OBS criterion), 435 436 Optimum damper size criterion (ODS criterion), 435 436 Optimum supplemental damping ratio, 339 340 Orthogonality condition, 50 OTM direction. See Overturning moment direction (OTM direction) Outer hollow section, 192 Output-only modal estimation (OOME), 61 62 Outrigger connection, 121 122 damping, 121 122, 122f systems, 958 Outrigger trusses, 762 OVE/OVE devices. See Spring fluid viscous dampers Overstrength factor, 284 Overturning moment direction (OTM direction), 75 76

P Pacific Earthquake Engineering Research Center (PEER Center), 706 707 PEER/ATC 72 1 code, 71 tall building initiative, 69 71 Pall Dynamics, 195 Pall system, 171 172

friction damper hysteretic steps, 174f for moment-resisting frames, 173f Panel deformation, 18 Pangu Plaza, Beijing, China building overview, 567f comparison different design strategies cost benefit analysis, 576 damping solution consideration, 576 structural options consideration, 576 damper distribution, 570f damper parameters, 573t damping overview architectural integration strategy, 575 damping strategy, 572 damping type, 573 elevator and other mechanical devices integration strategy, 575 experimental tests, 575 maintenance strategy, 576 MEP integration strategy, 575 monitoring system, 576 structural and damping design, 574 575 economic comparison with different plans, 577t framing plan, 570f fundamental frequencies and mode shapes, 569f introduction/history, 568 lesson learned and recommendations, 576 577 original and new structure dynamic properties, 577t project data, 567 568 structural periods and mass participation factors, 569t structural system, 568 571 additional damping provided by damping system, 570 building code, 571 building cost vs. damping cost, 571 building fundamental periods, 568 damping strategy utilization, 569 design forces, 571 peer-reviewed project, 571 summary of damper vibration absorption effect, 578t 3d structural model, 574f time history wave diagrams, 571f, 573t Parallelogram-type hysteretic model, 310 Parametric analysis, 695 Park Tower, 199, 199f Passive control strategy, 453 Passive damping systems, 108 206, 264 347, 772. See also Isolation systems distributed damping approaches, 109 195 mass damping approaches, 195 206 passive added damping devices, 108t

Index 1091

risk category of tall buildings, 268t step-by-step procedure for distributed dampers, 265 321, 266f for isolated dampers, 321 347 Passive energy dissipation systems, 108 Passive response mode, 756 Passive systems, 8 9, 107 108 PBD. See Performance-based design (PBD) PBSD. See Performance-based seismic design (PBSD) P-Delta analysis, 260 effects, 18, 84, 363 Peak Ground Acceleration (PGA), 18, 772 Peak modal responses, 253 PEER Center. See Pacific Earthquake Engineering Research Center (PEER Center) Peer-reviewed project, 738 PEM. See Pseudo-excitation method (PEM) Pendulum, 814 isolators, 216 217 TMD, 198 200, 199f, 1004 1005 Penguin vibration damper (PVD), 195 Performance level, 18 Performance objectives (PO), 244, 405 410, 406t seismic hazard level, 406 407 target building performance level, 407 410, 408f Performance testing, 474 478 Performance-based design (PBD), 1, 18, 87, 262 approach, 251 foundation, 2f limit states, 2f Performance-based earthquake resistant design, 87 89 Performance-based seismic design (PBSD), 703 704 Period of isolated system, 351 Periodic excitation. See Shock excitation PFF. See Pin Fuse Frame (PFF) PGA. See Peak Ground Acceleration (PGA) Phase transformation, 151 Piecewise linear models, 62 63 Piezoelectric dampers Application, 927 Pin Fuse Frame (PFF), 610 611 Pinnacle first approach, 766 TMDs, 767 768, 768f, 771 772 Piston diameter, 390 391 Plan distribution, 300 Plane-defining transition, 16 Plastic hinge region, 18 Plastic Wen model, 314 315, 315f PO. See Performance objectives (PO)

Polytetrafluoroethylene disk (PTFE disk), 214, 216 Postextreme event inspections, 519 524 base isolation systems, 523 524 distributed damping systems, 521 522 isolated damping systems, 522 Posttensioned energy dissipating system (PTED system), 182 183, 183f Posttensioned rocking systems, 180 183, 183f Posttensioning techniques, 18 Postyield stiffness, 358 Potential flow theory, 202 203 Power dissipation, 133 law models, 62 63 Power spectral density (PSD), 693 Practical design aspects, 248 251 damper system from concept to production process, 248 249 damping devices properties, 251 seismic design process and strength requirements, 250 251 wind design and occupant comfort process, 249 250 Pre-installation test, 18 active, semiactive and hybrid damping systems, 492 493 base isolation systems, 486 491 distributed damping systems, 473 482, 491 492 isolated damping systems, 492 isolated passive, active, semiactive and hybrid damping systems, 482 486 and quality control, 472 493 Precontrol principles, 526 Preliminary analyses, 260 261 Preliminary design, 899 Preliminary structural analysis, 354 357 Pressure relief valves (PRVs), 130, 135 136 Prestressed concrete, 18 Pretensioning techniques, 18 Primary building characteristics, 284 Primary frame system, 983 Principal inspection, 518 Probabilistic risk-targeted maximum considered earthquake ground motions, 273 Probabilistic seismic hazard analysis (PSHA), 101 Production test, 18 Production test, 479 Proof Pressure Test, 635 Property modifications factors, 263 264 Proportionality factor, 310 Prototype test(ing), 18, 468 469, 488 489 Provision Update Committee (PUC), 246 247 PRVs. See Pressure relief valves (PRVs)

1092 Index

PSD. See Power spectral density (PSD) Pseudo-excitation method (PEM), 1012 PSHA. See Probabilistic seismic hazard analysis (PSHA) PTED system. See Posttensioned energy dissipating system (PTED system) PTFE disk. See Polytetrafluoroethylene disk (PTFE disk) PUC. See Provision Update Committee (PUC) “Pushover” analysis, 17 PVD. See Penguin vibration damper (PVD)

Q QAP. See Quality assurance plan (QAP) Qualification tests, 19, 489 Qualification/production testing, 469 Quality assurance, 364 365 and experimental evaluation, 346 347 Quality assurance plan (QAP), 421 422 Quality control, 19 maintenance, and inspection requirements, 321 procedures, 490 Quasi-optimal distribution, 434 Quasi-static test, 19, 484, 490 Quintuple pendulum system principles and force displacement behavior, 216f

R Racking deformation, 19 Radiation damping, 14 Raffles city, Chongqing, China, 819f bearings location, 821f comparison different design strategies cost benefit analysis, 825 826 damping solution consideration, 825 structural options consideration, 824 825 complete LS-DYNA analysis mode, 824f conservatory/tower interface, 825f damping articulation scheme, 822f damping overview architectural integration strategy, 823 elevator and mechanical devices integration strategy, 824 experimental tests, 824 MEP integration strategy, 823 monitoring system, 824 strategy, 822 structural and damping design, 822 823 type, 822 fundamental frequencies and mode shapes, 820f introduction/history, 819 lesson learned and recommendations, 826 827 project data, 818 819

structural system, 819 822 additional damping provided by damping system, 821 building code, 821 822 building cost vs. damping cost, 821 building fundamental periods, 820 damping strategy utilization, 821 peer-reviewed project, 822 viscous dampers design parameters, 827t Ramberg Osgood model, 143, 144f Random, 30 31 Random Decrement technique (RD technique), 500 Random excitation, 397, 991 TMDs SDOF simplified procedure for, 995 999 Rating tiers, 527 528 Rayleigh damping, 14, 55 56, 413 Rayleigh-type damping systems, 434 435 RB models. See Replacement Beam models (RB models) RBS. See Reduced Beam Section (RBS) RC. See Reinforced concrete (RC) RD technique. See Random Decrement technique (RD technique) Real-time hybrid testing method (RTHTM), 492 Recentering spring systems, 178 180 Rectangular tanks, 336 339 REDi rating systems. See Resilience-based earthquake design initiative rating systems (REDi rating systems) Reduced Beam Section (RBS), 593 Redundancy, 19 Regular inspection, 518, 838 839, 845 846, 853, 861 Reinforced concrete (RC), 243, 602 buildings, 124 125 code/guideline intrinsic damping comparison for RC buildings, 74f structures, 66 Relative displacement, 125 126 Reliability factor (SF), 131 Replacement Beam models (RB models), 260 261 Resetting semiactive stiffness dampers (RSASD), 228 229 Residential functions, 958 961 Resilience, 19, 103 Resilience-based earthquake design initiative rating systems (REDi rating systems), 11 12, 526 531, 528f, 528t criteria, planning, and assessment, 527 528 and response control systems, 528 529 Resilient buildings, 9

Index 1093

Resistance factor, 19 Response acceptability checking, 317 321, 346 acceleration criteria, 319 320 damper criteria, 320 321 design review of damping systems and testing programs, 321 drift criteria, 319 structural system, 319 Response modification coefficient, 284 Response spectrum, 19, 268 275 analysis, 359 360 procedure, 289 297 Response-history analyses (RHA), 238 Response-spectrum method, 288 289 Response history procedure, 19 Responses, determination of, 397 Restoring force, 126 127, 135 136, 180 Retrofit, 19 strategies, 417 419 configuration and distribution for dynamic modification system, 418 select total target damping, 418 419 selecting suitable type of dynamic modification devices, 418 Retrofit of existing buildings, 400 429 code requirements, 400 402 evaluation procedures based on ASCE 41 13, 402 403 high-rise building 24-story building in Osaka with oil dampers, 425 30-story building in Osaka with viscous dampers, 425 428 35-story building in San Francisco with viscous dampers, 422 425 for building in Osaka TMD, 428 429 step-by-step procedure, 403 422, 404f acceptance criteria, 414 417 checking retrofit response acceptability, 420 421 damping system preliminary design, 419 evaluation of existing building, 410 412 initial considerations, 403 405 model and analyze existing building, 413 414 performance objectives and hazard levels, 405 410 quality control, maintenance, and inspection requirements, 421 422 retrofit strategies, 417 419 update building model and perform analyses, 419 420 Return period, 19

RFD device. See Rotation friction damper device (RFD device) RHA. See Response-history analyses (RHA) Rhombic damper, 154 155 configuration, 155f Riccati’s equation, 366 369 Risk category, 19, 268 risk-targeted maximum considered response spectrum, 272 Risk coefficient (CR), 273 Risk-targeted maximum considered earthquake (MCER), 269 274 roof displacement in modal forms, 295 Ritz-Vectors, 259 260 RMS. See Root-mean-square (RMS) Robinson Seismic Limited, 195 Robustness performance, 102 103 Rocker pendulum, 198 Roof, 454 Root-mean-square (RMS), 204, 510, 742 Rotating equipment, 32 Rotation friction damper device (RFD device), 172 174 4 joint type, 174f with multiunits, 175f Rotational friction system, 172 176 frame structure, 175f 2000 kN damper installing in tallest building, 177f Round bars absorbers, 164 Rowan Williams Davies & Irwin, Inc. (RWDI), 815 RSASD. See Resetting semiactive stiffness dampers (RSASD) RTHTM. See Real-time hybrid testing method (RTHTM) Rubber, 141 bearings, 462 isolator, 244 245 compression stress limits, 244t and sliding bearings, 524 Rule of thumb, 207 208 RWDI. See Rowan Williams Davies & Irwin, Inc. (RWDI)

S Safety factor, 131 Safety gear systems, 451 453 San Diego Central Courthouse, San Diego, California, USA building overview, 592f comparison different design strategies

1094 Index

San Diego Central Courthouse, San Diego, California, USA (Continued) cost benefit analysis, 614 damping solution consideration, 611 613 structural options consideration, 610 611 damping overview architectural integration strategy, 605 606 damping strategy, 597 damping type, 597 598 elevator and other mechanical devices integration strategy, 606 experimental tests, 606 608 maintenance strategy, 610 MEP integration strategy, 606 monitoring system, 608 610 structural and damping design, 599 605 damping ratio formula, 598f, 599f fundamental frequencies and mode shapes, 595f introduction/history, 593 lesson learned and recommendations, 614 615 project data, 592 593 seismic design criteria summary, 603t structural system, 593 597 additional damping provided by damping system, 595 597 building code, 597 building cost vs. damping cost, 597 building fundamental periods, 593 damping strategy utilization, 593 595 peer-reviewed project, 597 typical transverse frame elevation, 594f typical upper level framing plan, 594f San Francisco with viscous dampers distribution schemes for damping constant, 423f plan and exterior frames, 423f retrofit strategies for 35-story building in, 422 425 Sandwich beam (SWB), 260 261 SAP2000, 314, 345 SAPHIA damper, 180, 182f SASDs. See Semiactive stiffness dampers (SASDs) SATMD. See Semiactive tuned mass dampers (SATMD) SB. See Shear beam (SB) SB system. See Sliding bearing system (SB system) SBB. See Seismic ball bearings (SBB) Scaled building model, 834 835, 859 Scissor, 118 connection damper system, 119f Scorpion yielding connector (SYC), 142, 166 167, 166f Scorpion yielding system, 194 Screening procedure. See Tier 1 procedure

Scruton helix, 34 SDOF. See Single degree of freedom (SDOF) Seal(s), 131 134 designs, 132 life, 132 materials, 131 132 SEAONC. See Structural Engineers Association of Northern California (SEAONC) SEAVANS twin building, 124 Seismic activity, 501 application, 130 base isolation, 210 base shear, 291 control, 30 control of industrial machine structures, 32 33 bearing supported fans in air conditioning system and power plant, 32f elastic support of diesel engine principle, 33f turbine generator machine buildings, 33f vibration isolation of transformer and with isolators, 34f deflection criteria, 90 design category, 270, 272t process and strength requirements, 250 251 requirements, 466 467 events, 98 forces, 57 hazard level, 20, 406 407 isolation layer, 449 450 system, 107 108, 351 352, 895, 900 loading, 17, 226 performance factors, 88 response damping role in seismic response control, 89 spectrum, 19 seismically isolated buildings, 456 Seismic ball bearings (SBB), 214 215 Seismic force resisting system (SFRS), 19, 601 602 Seismic loading floor accelerations under, 89 hysteretic damping, 124 125 inelastic behavior, 288, 327, 369 site-specific study, 707 Seismic Rehabilitation, 400 Seismic Rehabilitation of Existing Buildings, 247 Seismic Retrofit, 400 Seismic testing. See Velocity performance testing Self-centering system(s), 143, 176 183 elastoplastic and self-centering structure, 177f frequency response curve, 178f, 179f

Index 1095

idealized displacement, 145f posttensioned rocking systems, 180 183 recentering spring systems, 178 180 Self-Mass Damper (SMD), 450 Semiactive, active, and hybrid dampers manufactures, 232 Semiactive control, 226 of base isolation systems, 225 with controllable hydraulic device, 226f devices, 226 227 semiactive controllable fluid dampers, 227 228 strategies, 226 227 systems, 219 220, 221f, 518 Semiactive dampers, 397 399 Semiactive damping systems, 492 493, 497t Semiactive devices, 226 227 Semiactive resetable device, hysteresis behavior of, 394f Semiactive stiffness dampers (SASDs), 228 229 Semiactive systems, 217 232, 499 500, 517 518, 926 active mass damper, 510 active, semiactive, hybrid dampers manufactures, 232 actuators, 223 224 ATMD, 222 223, 229 230, 510 512, 511f case studies, 534, 904 919 Thyssenkrup Test Tower, Rottweil, Germany, 904 919 control strategies, 230 231 dampers manufactures, 232 design, 383 394 future directions, 231 232 HMD, 512 513 interaction of active damping system and DSF systems, 460f literature review, 365 367 SASDs, 228 229 SATMDs, 226 227 semiactive controllable fluid dampers, 227 228 semiactive, hybrid dampers manufactures, 232 step-by-step procedure, 367 400, 368f bandwidth specification of active and semiactive systems, 370 build simplified and reduced model, 372 376 building and site categorization, 369 building fundamental properties and preliminary structural analyses, 369 370 damper type, configuration, and distribution of sensors and actuators, 370 372 design of controller system, 372 376 selecting structural system, 369 selection of performance objective, 370

verification analyses and quality control, 394 400 Semiactive tuned mass dampers (SATMD), 226 227, 391, 399 400 design, 391 394 model, 392f strategy, 227f Semiactive viscous dampers, 399 design, 387 391, 388f design parameters of 2 4 semiactive damper, 391f hysteresis for 1 4 device, 1 3 device, and 2 4 device, 388f hysteretic loops of 2 4 device and 1 4 device, 389f Sendai MT Building, 210 Sensors, 517 Service-level earthquake (SLE), 69 71, 268 269 Serviceability, 17 Serviceability Limit State (SLS), 79, 246 Servo-hydraulic cylinders, 198 SFRS. See Seismic force resisting system (SFRS) Shaking table hybrid testing, 486 test, 20, 481, 485, 490 491 Shallow water TLD, 201 202 wave theory, 202 203 Shanghai Tower, Shanghai, China, 199, 445, 446f, 864f acceleration vs. return period, 869f comparison different design strategies cost benefit analysis, 875 damping solution consideration, 875 structural options consideration, 874 875 damping overview, 869 874 architectural integration strategy, 873 elevator and mechanical devices integration strategy, 874 experimental tests, 874 maintenance strategy, 874 MEP integration strategy, 873 monitoring system, 874 strategy, 870 structural and damping design, 873 type, 871 872 fundamental frequencies and mode shapes, 868f introduction/history, 864 865 lesson learned and recommendations design innovative solutions, 875 876 difficulties in design, 875 improvements, 876 peak acceleration reduction, 871t

1096 Index

Shanghai Tower, Shanghai, China (Continued) project data, 863 structural system, 865 868 additional damping provided by damping system, 868 building code, 868 building cost vs. damping cost, 868 building fundamental periods, 866 damping strategy utilization, 866 868 design forces, 868 expected performance, 868 peer-reviewed project, 868 TMD general scheme, 870f typical elevation, 867f typical MEP/refugee structural framing, 866f typical structural framing plans, 865f Shape memory alloy damper, 14 Shape Memory Alloy Devices (SMAD), 194 195 Shape memory alloys (SMA), 146, 151, 927 behavior, 151, 151f Shear, 20 deformations, 111 115 force, 357 shear-frame systems, 958 shear-type model, 20, 244 strain compatibility, 976 wall, 20 Shear beam (SB), 260 261 Shear panel damper (SPD), 184 186, 184f fish plate connecting infill plates and frame members, 186f with rib plates, 185f steel shear wall with slits, 186f and timber panel configuration, 187f Shinjuku Center Building in Tokyo, 9 Shiodome Sumitomo Building, 208 209, 448 449 Shizimu Corporation, 210 SHM. See Structural health monitoring (SHM) Shock absorbers, 811 Shock excitation, 31 Shock isolation system, 178 Shoulder 1 Apex dampers (TMD or TLD), 699 700 Silicon-based elastomer, 180 Simple-pendulum TMD, 765 766 Single concave friction pendulum, 360 361 Single degree of freedom (SDOF), 19, 239 240, 240f, 491 system, 20, 41 47, 41f, 111 115, 375f, 377 379, 977 978 equations of motion, 41 42

forced vibration properties, 43 47 free vibration properties, 42 43 optimum TMD parameter for damped structure modeled, 989t optimum TMD parameter for undamped structure modeled, 986t Single mass damper, 329 Single round-hole damper, 156 Single TLCDs, 339 340, 343 344 Single TMDs, 333 Single-acting device, 180 “Single” isolators, 214 Sinusoidal Performance Test, 635 Site class, 20, 269 270, 270t Site coefficients, 270, 271t Site material and equipment protection, 526 Site soil condition, 350 Site spectral response acceleration, 268 275 Site-specific response spectrum, 251 Skeleton curve, 146, 147f Skilling Ward Rogers Barkshire (SWRB), 539, 547 SLE. See Service-level earthquake (SLE) Slender buildings, 7 8 Slenderness ratio, 20 Sliding, 210 mode controllers, 231 seal, 132 133 systems, 213 215 Sliding bearing system (SB system), 210, 351 352 Slip force of friction damper, 311 Slit shear steel plates, 186 Sloshing dampers, 200 201 liquid, 203 Slotted-bolted connections, 171 steel frame, 172f SLS. See Serviceability Limit State (SLS) SMA. See Shape memory alloys (SMA) SMAD. See Shape Memory Alloy Devices (SMAD) “Smart” fluids, 227 228 “Smart” systems. See Semiactive control systems SMD. See Self-Mass Damper (SMD) SMF. See Special Moment Frame (SMF) Smoothed particle hydrodymanics (SPH), 202 203 S-N diagram. See Stress-Number of cycles diagram (S-N diagram) Snubber ring, TMD, 770 771, 771f Snubber system, 872

Index 1097

Softened dynamic seal, 133 Software, 438 Soil structure interaction (SSI), 20, 261 262 SPAF. See System property adjustment factor (SPAF) SPD. See Shear panel damper (SPD) Special connecting devices, 142, 169 183 friction dampers, 169 176 self-centering systems, 176 183 using weight of construction, 183 Special Moment Frame (SMF), 593, 610 611 Special structural members, 142, 184 192 BRBs, 186 191 SPD, 184 186 TTD, 192 Spectral, 19 acceleration effect of building stiffness, 89f effect of damping, 90f displacement, 273 274 Spectrum matching, 274 275 SPH. See Smoothed particle hydrodymanics (SPH) Spillover phenomenon, 376 Spring fluid viscous dampers, 129, 134 Spring force. See Restoring force Spring(s), 135 136 spring-friction damper system, 310, 310f spring/damper assembly, 118 stiffness and mass, 834 835 Square root of Sum of Squares (SRSS), 293 294 SRC. See Steel RC (SRC) SRSS. See Square root of Sum of Squares (SRSS) SSASD. See Switching semiactive stiffness dampers (SSASD) SSD. See Steel slit damper (SSD) SSI. See Soil structure interaction (SSI) St Francis Shangri La Place, Manila, Philippines, 558f, 560f building overview, 557f close-up of damper, 563f comparison different design strategies cost benefit analysis, 566 damping solution consideration, 566 structural options consideration, 566 damper installation, 564f damping overview architectural integration strategy, 565 damping strategy, 561 damping type, 561 563 elevator and other mechanical devices integration strategy, 565 experimental tests, 566 maintenance strategy, 566 MEP integration strategy, 565

monitoring system, 566 structural and damping design, 563 565 introduction/history, 557 lesson learned and recommendations, 566 outrigger damper, 562f project data, 556 557 structural system, 557 561 additional damping provided by damping system, 559 560 building code, 561 building cost vs. damping cost, 561 building fundamental periods, 558 damping strategy utilization, 559 design forces, 561 expected performance, 561 peer-reviewed project, 561 variation in lateral accelerations, 565t in wind overturning moment with damping, 565t Stage 1 ETABS v9.7 model, 602 Standard deviation of roof displacement, 395 396 Standard devices, 142, 152 169 ADAS, 152 153 CPD, 160 DFMDs, 155 156 HDS, 155 KBF, 164 166 LEDs, 167 169 rhombic damper, 154 155 SSD, 156 160 SYCs, 166 167 TADAS, 154 U-shaped metallic damper, 160 163 yielding steel bracing system, 164 Static linear analysis, 237 238 Statistical assessment, 394 Steady-state harmonic oscillation, 53 Steel, 146 148, 953 behavior, 146 148 buildings, 75f cyclic stress strain hysteresis, 146f frame, 735 hysteresis loop decomposition, 147f plates, 212 shear wall connection, 185 structures, 66, 572, 585 damage ratios drifts, 99t damage ratios floor acceleration, 100t Steel damper. See U-shape damper Steel Outrigger Trusses, 865 Steel RC (SRC), 819 820 Steel slit damper (SSD), 152, 156 160 configuration, 159f

1098 Index

Stick-slip behavior, 61 stick-slip model intrinsic damping in buildings, 61f “Stiction” model, 61 Stiffness stiffness-proportional damping distribution, 305, 308 equation, 55 of TMDs, 333 334 Stochastic wind-force model, 693 Storage/installation of damping systems, 525 526 Story, 20 Story displacement. See Relative displacement Story drifts, 84 85, 319t, 620 Story shear-proportional damping distribution, 309 “Straight one” shape, 718 Strain, 145 independency, 137 Strain versus Cycles diagram. See Stress-Number of cycles diagram (S-N diagram) Strength design, 20 basic load combinations for, 283 Stress-Number of cycles diagram (S-N diagram), 502, 504 Stress strain curve, 145 for aluminum alloy and steel, 150f Strip damper, 156 Structural analyses, 251 264, 312 317 boundary conditions and common assumptions, 261 264 computational structural analysis modeling, 252 damper synthesis of RB idealization process, 261f effective stiffnesses of RC structural elements, 256t, 257t, 258t expected material strengths, 255t linear analysis methods, 252 255 linear methods for supplemental damping systems, 253 255 nonlinear analysis methods, 255 260 preliminary analyses, 260 261 type, 252 261 Structural behavior improved model, 922 925 enhanced computing power, 923 924 improved design codes, 924 925 improving behavior model of materials and devices, 923 improving understanding of actions on structures, 922 923 Structural control and health monitoring, 102 103 Structural engineering, 3 4 Structural Engineers Association of Northern California (SEAONC), 246 247 Structural health monitoring (SHM), 505 506

Structural material, 953 957 tallest buildings according to CTBUH skyscraper center, 955t used in buildings with damping system, 960f Structural reliability, 1 2 Structural system, 1, 3, 20, 312 313, 319, 958 selection, 349 351 used in buildings with damping system, 960f Stud-type systems, 439, 440f Sub-structures online test, 18 Substructure hybrid test, 20, 486 Substructured online test, 491 Sumitomo friction devices, 171, 172f Sumitomo Riko Company Limited, 141 Superelastic. See Austenite Supertall buildings, 949 950 Supervisory control scheme, 232 Supplemental damping system, 209, 251 linear methods for, 253 255 Supplementary damping, 9 10, 76, 949 systems, 30 Surface-defining transition, 16 Swatch Building, 450, 452f SWB. See Sandwich beam (SWB) Switching semiactive stiffness dampers (SSASD), 228 229 SWRB. See Skilling Ward Rogers Barkshire (SWRB) SYC. See Scorpion yielding connector (SYC) System characterization tests, 487 System property adjustment factor (SPAF), 264

T TADAS systems. See Triangular added damping added stiffness systems (TADAS systems) Taipei 101, Taipei, Taiwan, 199, 445, 446f, 758f, 760f, 761f, 764f ball-shaped tuned mass damper, 766f base bending moment induced by vortex induced oscillation, 769f comparison different design strategies cost benefit analysis, 774 damping solution consideration, 773 structural options consideration, 773 damping overview, 765 773 architectural integration strategy, 772 773 elevator and mechanical devices integration strategy, 773 experimental tests, 773 MEP integration strategy, 773 monitoring system, 773 strategy, 765 766

Index 1099

structural and damping design, 768 772 type, 767 768 introduction/history, 759 lesson learned and recommendations design innovative solutions, 774 difficulties in design, 774 improvements, 774 project data, 758 759 structural system, 759 765 additional damping provided by damping system, 764 building code, 764 765 building cost vs. damping cost, 764 building fundamental periods, 762 damping strategy utilization, 762 764 design forces, 765 peer-reviewed project, 765 Tall building with damping systems database 1 (worldwide buildings over 250 m) general trends for tall buildings, 930 general trends for tall buildings with damping systems, 930 937 trends for damping system category, 937 941 trends for damping system types, 941 943 database 2 (buildings over 200 m in United States) general trends for buildings, 943 945 general trends for damping system, 945 947 trends for damping system category, 947 trends for damping system type, 948 further studies, 948 961 building function, 958 961 damping vs. height, 949 953 structural material, 953 957 structural system, 958 summarized data of buildings with damper, 961 Tall buildings, 7 8, 208, 445, 513 515, 949 953 comparison of dampers, 232 damping trends, 954f dynamic modification devices, 28 30 earthquake-excited motion, 85 90 damping role in seismic response control, 89 principles of performance-based earthquake resistant design, 87 89 seismic deflection criteria, 90 earthquake-excited motion, 85 90 damping role in seismic response control, 89 performance-based earthquake resistant design, 87 89 seismic deflection criteria, 90 possible supplementary damping systems for, 8f wind-excited motion, 78 85

building wind vibration, 79 80 occupant comfort, 80 84 wind deflection criteria, 84 85 Tall Buildings Initiative (TBI), 268 269, 706 707 Tank length, 335 Tapered-beam dampers, 141 142 Target vibration period, 351 Taylor Devices, 29, 135 136, 195, 794 795 Taylor Fluid Viscous Dampers, 135 136 Taylor viscous dampers, 635 636 TB. See Timoshenko beam (TB) TBD. See Toggle-brace damper (TBD) TBI. See Tall Buildings Initiative (TBI) Technical Subcommittee 12 (TS 12), 246 247 Technical University of Denmark (DTU), 492 Technology transfer, 1 Temperature dependency, 975 Tensa International, 217 Tension-compression isolator, 178 TESolution, 206, 232 Testing, 11 12 Testing, inspection, and maintenance. See also Damping system(s) building health monitoring, 505 513 codes and standards development, 466 472 Chinese code, 471 European standards, 472 Japanese code, 470 471 US standards, 466 470 commissioning and system tuning, 493 500 fatigue of dampers, 501 505 maintenance cost, 519 maintenance issues for different control systems, 520t ongoing maintenance, 513 519 active, semiactive and hybrid damping systems, 517 518 base isolation systems, 518 519 distributed damping systems, 516 isolated damping systems, 516 517 standard recommendations, 515 postextreme event inspections, 519 524 base isolation systems, 523 524 distributed damping systems, 521 522 isolated damping systems, 522 preinstallation tests and quality control, 472 493 base isolation systems, 486 491 distributed damping systems, 473 482 isolated passive, active, semiactive and hybrid damping systems, 482 486 testing examples, 491 493 REDi rating systems, 526 531

1100 Index

Testing, inspection, and maintenance (Continued) transportation, installation and care of works, 524 526 USRC rating systems, 526 531 Thermodynamics, 133 Thermoplastic rubber, 124 36-story building, 512 Thousand tower, 209 Three modes of building integration, 438 Three-dimensional dynamic properties (3D dynamic properties), 252 Three-DOF model, 376 3M, 140 141 3M ISD, 503 3M ISD-111 material 3M ISD-111H material, 138 139 series of material, 136 137 3M VE dampers, 645 material, 503 Thyssenkrup Test Tower, Rottweil, Germany, 905f comparison different design strategies cost benefit analysis, 918 damping solution consideration, 917 918 structural options consideration, 917 damping overview architectural integration strategy, 914 control strategies, 916 917 damping strategy, 909 elevator and other mechanical devices integration strategy, 914 experimental tests, 914 916 MEP integration strategy, 914 monitoring system, 916 structural and damping design, 909 913 type, 909 feedback control loop practice, 916f fundamental frequencies and mode shapes, 908f introduction/history, 905 906 lesson learned and recommendations, 918 919 design innovative solutions, 919 difficulties in design, 918 919 possible improvements, 919 pendulum rope supported TMD mass, 910f project data, 904 905 structural system, 906 909, 910f additional damping, 908 building code, 909 building cost vs. damping cost, 909 building fundamental periods, 907 damping strategy, 908 peer-reviewed project, 909 plan view, 906f structural elevation, 907f wind tunnel test model, 911f

Tianjin International Trade Center, Tianjin, China, 116 117 building overview, 626f comparison different design strategies cost benefit analysis, 636 damping solution consideration, 636 structural options consideration, 636 damper installation, 631f damping overview architectural integration strategy, 635 damping strategy, 631 damping type, 631 elevator and other mechanical devices integration strategy, 635 experimental tests, 635 maintenance strategy, 635 MEP integration strategy, 635 monitoring system, 635 structural and damping design, 632 634 elevation picture, 630f introduction/history, 627 lesson learned and recommendations design innovative solutions, 637 difficulties in design, 637 improvements, 637 project data, 626 627 structural system, 627 631 additional damping provided by damping system, 628 building code, 630 631 building cost vs. damping cost, 628 629 building fundamental periods, 627, 628f damping strategy utilization, 627 peer-reviewed project, 631 tower retrofit scheme, 633f Tier 1 procedure, 402 Tier 2 procedure, 402 Tier 3 procedure, 403 Time histories, 268 275 analyses, 244, 393 394 Time-dependent element. See Dynamic element Timoshenko beam (TB), 260 261 TJBDPA. See Japanese Building Disaster Prevention Association (TJBDPA) TLCD. See Tuned liquid column damper (TLCD) TLD with Floating Roof (TLD-FR), 336 TLD-FR. See TLD with Floating Roof (TLD-FR) TLDs. See Tuned liquid dampers (TLDs) TLMD, 200 TMDs. See Tuned mass dampers (TMDs) Toggle frames, 116 117 dampers in toggle-braced frames, 117f toggle connection damper system, 118f types of toggle configuration, 117f Toggle-brace damper (TBD), 116 117, 634

Index 1101

Tokio Skystree east tower, 208 Topographic factor, 276 Torsional steel “springs”, 742 Total damping, 14 ratio selection, 351 Total design displacement, 15 Total maximum displacement, 15 Total seismic base shear, 293 294 Total target damping selection, 298, 327 328 Tower, 803 and conservatory models, 823 Transaction rating, 530 Transfer systems, 353 Transient, 30 31 Translation TMD, 197 198, 198f TRANSPEC SHA, 134 Transportation of damping systems, 524 525 Transverse direction, 748 Triangular added damping added stiffness systems (TADAS systems), 154 155 device hysteretic loop and dimensional parameters, 154f Trigonometry, 976 977 “Triple” isolators, 214 pendulum system principles and force displacement behavior, 216f Truss system, 449 450 TS 12. See Technical Subcommittee 12 (TS 12) TSD. See Tuned-slosh damper (TSD) TSDs. See Tuned sloshing dampers (TSDs) TTD. See Tube-in-tube damper (TTD) Tube dimensions, 343 344 Tube systems, 958 Tube width, 343 344 to liquid length ratio, 343 344 Tube-in-tube damper (TTD), 142, 192 typical damper configuration and equivalent strip geometry, 194f Tuned liquid column damper (TLCD), 14, 204 206, 204f, 235 236, 323 324, 446, 484 485, 499, 785, 785f, 948, 1008 1009 bidirectional, 205f comparison of optimal parameters TMD and, 1010t design parameter determination, 339 344 equivalent mechanical properties, 344 mass ratio, 332 modeling, 345 346 optimality of TLCD parameters, 436 schematic representation of SDOF with, 1009f system, 457 Tuned liquid dampers (TLDs), 14, 37, 200 206, 235 236, 322 323, 446, 484, 943, 1005 1007

analytical and numerical fluid models, 202 203 determination of design parameters, 334 339 equivalent mechanical models, 203 204 equivalent mechanical properties, 336 339 mass ratio, 332 modeling, 345 compared to SDOF system, 1005f performance charts of equivalent linear TLD, 1006f rectangular, 202f schematic view of modified-TLD-structure model, 1008f tank design, 1007 tank dimensions, 335 TLCD, 204 206 and TMD dual damper experimental setup, 201f Tuned mass dampers (TMDs), 14, 37, 78, 109, 195 200, 225, 235 236, 254, 321 322, 483 484, 495 498, 519, 566, 735, 763, 763f, 870, 871f, 943, 948, 983 1005 advantages and disadvantages, 196t applications, 200 control system, 740 742 design parameters determination, 333 334 determination of mass, stiffness, and damping coefficients, 333 334 mass ratio, 332 MDOFs system, 999f theory, 999 1001 modeling, 345 MTMDs theory, 1001 1003, 1002f optimality of TMD tuning ratio, 436 optimum damping ratio, 991f optimum TMD parameter for damped structure modeled as SDOF system, 989t for undamped structure modeled as SDOF system, 986t optimum tuning frequency, 991f pendulum, 198 200, 445 446, 1004 1005 properties, 983, 993 retrofit strategies for 27-story building in Osaka, 428 429 SDOF simplified procedure, 993 999 theory, 983 992, 984f system, 738 and TLD dual damper experimental setup, 201f TMD-position-dependent nonlinear damping, 872 TMD/TLD/TLCD manufactures, 206 TMD DFS damping interaction, 458 459 translation, 197 198

1102 Index

Tuned mass dampers (TMDs) (Continued) translational/sliding, 446 Tuned Mass Inerter systems, 928 Tuned sloshing dampers (TSDs), 200 201, 254 system, 457 plan adjacent to mechanical room located at center, 457f Tuned-mass system, 132 Tuned-slosh damper (TSD), 566 Tuning bandwidth frequency, 342t Tuning frequency, 226, 322, 339 340, 1001 1002 ratio, 196 Tuning ratio, 203, 1005 Turbulence, 8 9 Turkey code, 73 20-story based isolated building, 513 Twisting-beam dampers, 141 142 Two Union Square, Seattle, USA, 555f building overview, 548f comparison different design strategies, 555 556 cost benefit analysis, 556 damping solution consideration, 556 structural options consideration, 556 damper mid installation, 554f damping overview architectural integration strategy, 553 damping strategy, 551 552 damping type, 552 elevator and other mechanical devices integration strategy, 554 experimental tests, 554 maintenance strategy, 555 MEP integration strategy, 554 monitoring system, 555 structural and damping design, 552 553 introduction/history, 547 548 lateral system, 550f lesson learned and recommendations design innovative solutions, 556 difficulties in design, 556 possible improvements, 556 project data, 547 structural system, 548 551 additional damping provided by damping system, 551 building code, 551 building cost vs. damping cost, 551 building fundamental periods, 549 damping strategy utilization, 549 551 peer-reviewed project, 551 3d model of damper, 553f tower framing plan, 549f viscoelastic damper, 552f

Two-DOF system, 375 Type test (European Standard). See Prototype test (ing)

U UAE. See United Arab Emirates (UAE) ULS. See Ultimate Limit State (ULS) Ultimate, 17 Ultimate earthquake loads, 692 Ultimate Limit State (ULS), 79, 246 wind loads, 691 Unacceptable response, 317 318 Undamped structure, 984 Uniform damping distribution, 303 305, 307 309 Uniform-moment dampers, 141 142 United Arab Emirates (UAE), 937 United States (US), 210 codes, 246 247 damping device technology in, 466 standards, 466 470 design review, 469 inspection, periodic testing, and maintenance program, 469 470 prototype tests, 468 469 qualification/production tests, 469 testing, 473 474 United States resiliency council rating systems (USRC rating systems), 11 12, 526 531 components, 530t University of Western Ontario (UWO), 776 Unloading stiffness, 311 312 Upper bound period, 274 US. See United States (US) U-shape damper, 14, 142, 194 195, 473 478, 492, 505, 523 524 damage to, 524f testing processes for, 496t U-shaped metallic damper, 160 163 biaxial U-shaped metallic-yielding damper, 162f frame configuration, 163f geometry, 162f uniaxial U-shaped metallic-yielding damper, 161f US Panel on Structural Control Research (USNSF), 29 30 USRC rating systems. See United States resiliency council rating systems (USRC rating systems) UWO. See University of Western Ontario (UWO)

V V-brace system, 160 V-bracing, 187 188

Index 1103

VCDs. See Viscoelastic coupling dampers (VCDs) VDDs. See Viscous damping devices (VDDs) VDs. See Viscous dampers (VDs) VE. See Viscoelastic (VE) Velocity activated supplemental damping system, 254 dependent devices, 245 246 exponent, 302 feedback, 381 systems, 231 performance testing, 606 608 pressure exposure coefficients, 279 280, 280t Velocity-dependent dampers, 108 109 damping devices. See Viscous damping devices (VDDs) devices, 479 force, 124 125, 129 VE dampers, 126 141 viscous dampers, 126 141 Verified rating, 531 Vertex displacement, 624 Vertical distribution, 300 301, 330, 331f, 439 441 Vertical positions, 439 Vertical transportation (VT), 450 451 Vibration control, 30 31 of industrial machine structures, 32 33 elastic support of diesel engine principle, 33f turbine generator machine buildings, 33f vibration isolation of transformer and with isolators, 34f of large industrial chimney stacks, 34 37 ring-type tuned mass damper, 37f Vibration isolation, 31 33 of buildings and rail structures, 34 floating bridges on helical springs, 35f floating trackbed with 6 Hz spring support, 35f Vicoda, 206, 217 VIO. See Vortex-induced vibration (VIO) Viscoelastic (VE), 29 dampers, 14, 78, 108, 127f, 136 141, 443, 444f, 461, 473 478, 491 492, 503, 522, 540, 798, 975 978 with cables, 125 126 connected to steel beam and cables, 126f damper modeling description, 139 140 functional description, 137 138 high damping rubber, 139 hysteretic behavior, 977f installed in building, 544f 3M ISD-111H material, 138 139 material behavior, 138

property estimation, 306 308 specification requirements, 140 testing processes for, 495t Two Union Square, 552f type of devices and manufacturers, 140 141 typical scheme, 544f VE damping material, 136f viscoelastic device, 136f damping coefficient, 308 distributed damping system, 700 elements, 926 material, 121 122, 136 137 behavior, 137 layers, 124 125 manufacturers, 140 VE-plastic response, 652 654 wall dampers, 124, 124f Viscoelastic coupling dampers (VCDs), 121 122, 124 125, 141, 259, 439, 639, 645 embedded among RC walls, 440f outrigger, 123f VCD schematics, 710f Viscous dampers (VDs), 129, 473 476, 478, 491, 502 503, 516, 519, 521 522, 572, 585, 686, 941 properties estimation, 302 306 retrofit strategies for 30-story building in Osaka with, 425 428 enerdy-dissipation ratio vs. damping constant, 427f installed connecting dampers and zoomed details, 428f investigation, 427f main and extension building, 426f retrofit strategies for 35-story building in San Francisco with, 422 425 distribution schemes for damping constant, 423f plan and exterior frames, 423f testing processes for, 494t testing requirements, 131 Viscous damping devices (VDDs), 14, 20, 256 258, 593 597, 600 601, 610 611, 765 766, 769f, 926 Viscous fluid damper, 471 Viscous toggle-brace distributed damping system, 700 Viscous wall damper, 123f, 124, 134, 443 Viscous/viscoelastic dampers modeling, 314, 314f Vision 2000, 400 401 Vortex shedding, 34 resonant behavior, 79 80 Vortex-induced vibration (VIO), 766 VSL International Ltd., 141 VT. See Vertical transportation (VT)

1104 Index

W Wall, 20 configuration, 184 dampers, 122 124 Water depth, 335 Waterproofing, 323 Wave sensor, 807f 1151 West Georgia, Vancouver, Canada, 855f comparison different design strategies cost benefit analysis, 862 damping solution consideration, 862 structural options consideration, 862 damping overview, 857 862 architectural integration strategy, 860 development process, 861 862 elevator and mechanical devices integration strategy, 860 experimental tests, 860 861 maintenance strategy, 861 MEP integration strategy, 860 monitoring system, 861 strategy, 857 858 structural and damping design, 859 860 type, 858 859 fundamental frequencies and mode shapes, 856f introduction/history, 855 lesson learned and recommendations design innovative solutions, 863 difficulties in design, 862 improvements, 863 pendulum-type dynamic test rig, 860f preliminary sizing of TLCD options, 858f project data, 854 structural system, 855 857 additional damping provided by damping system, 856 building code, 857 building cost vs. damping cost, 856 building fundamental periods, 855 damping strategy utilization, 855 design forces, 857 expected performance, 857 peer-reviewed project, 857 wind tunnel test model, 856f Wind actions, 66 analysis, 69, 72 application, 130 deflection criteria, 84 85 demand, 275 283 design, 68 process, 249 250 directionality factor, 276, 277t excitation, 379 observed response under, 508

forces, 57 load, 17, 740 pressure, 281 restraint system, 20 tunnel data files, 786 procedure, 281 283 test, 20, 250, 481, 485, 790, 834 835, 835f turbines, 3 wind-induced bending moment, 796, 797f, 798f wind-induced vibration, 37 wind/seismic force demands, 3 Wind testing. See Cycle testing Worldwide buildings over 250 m (Database 1). See also Damping system(s) trends for tall buildings, 930, 931f trends for tall buildings with damping systems, 930 937 for countries of major regions in recent, 936f by decade, 932f distribution by region, 934f distribution of tall buildings, 933f by region, 934f in tall buildings, 931f in two time periods, 935f Worst-case analysis, 102 Wuhan Poly Cultural Plaza, Wuhan, China building overview, 616f comparison different design strategies cost benefit analysis, 625 damping solution consideration, 624 625 structural options consideration, 624 construction view, 619f damper parameter list, 622t damper site installation, 621f damping overview architectural integration strategy, 623 damping strategy, 620 damping type, 620 development process, 624 elevator and other mechanical devices integration strategy, 623 experimental tests, 623 maintenance strategy, 623 MEP integration strategy, 623 monitoring system, 623 structural and damping design, 621 622 introduction/history, 617 lesson learned and recommendations design innovative solutions, 625 difficulties in design, 625 improvements, 626 project data, 615 616 structural system, 617 620

Index 1105

additional damping provided by damping system, 617 618 building code, 619 building cost vs. damping cost, 618 building fundamental periods, 617, 618f damping strategy utilization, 617 design forces, 620 peer-reviewed project, 620 performance, 620

X X-shaped metallic damper, 155

Y YBS. See Yielding brace system (YBS) YC Condominiums, 639, 645 Yield stress, 20 Yielding damper category, 146 properties, 145 displacement, 310 311 limit, 150 metallic dampers, 522 ratio, 311 steel bracing system, 164 Tyler’s yielding steel bracing configuration, 164f stiffness, 311 312 Yielding brace system (YBS), 166 167 454 Yonge, Toronto, On, Canada building overview, 638f comparison different design strategies cost benefit analysis, 656 damping solution consideration, 654 656 structural options consideration, 654 damping overview architectural integration strategy, 651 damping strategy, 645 646 damping type, 646 elevator and other mechanical devices integration strategy, 651 experimental tests, 651 654 MEP integration strategy, 651 monitoring system, 654 structural and damping design, 646 651 decision factors by developer, 655t fundamental frequencies and mode shapes, 640f introduction/history, 639 lesson learned and recommendations, 656 658 design innovative solutions, 657

difficulties in design, 657 improvements, 658 plan view, 641f project data, 637 639 structural system, 639 644, 642f additional damping provided by damping system, 641 643 building code, 644 building cost vs. damping cost, 644 building fundamental periods, 640 damping strategy utilization, 640 641 peer-reviewed project, 644 wind tunnel test, 643f 14 York Street, Toronto, Canada, 840f, 841f comparison different design strategies cost benefit analysis, 846 damping solution consideration, 846 structural options consideration, 846 damping overview, 844 846 architectural integration strategy, 845 development process, 846 elevator and mechanical devices integration strategy, 845 experimental tests, 845 maintenance strategy, 845 846 MEP integration strategy, 845 monitoring system, 845 strategy, 844 structural and damping design, 844 845 type, 844 elevation, 842f introduction/history, 841 lesson learned and recommendations design innovative solutions, 846 difficulties in design, 846 improvements, 847 project data, 840 841 structural system, 841 843 additional damping provided by damping system, 843 building code, 843 building cost vs. damping cost, 843 building fundamental periods, 842 damping strategy utilization, 842 843 design forces, 843 expected performance, 843 peer-reviewed project, 843 Yozemi Tower, 209

Z Zero Leakage Technology, 134